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Low-cycle fatigue behavior of a cast Mg–Y–Nd–Zr alloy by T6 heat treatment
Materials Science & Engineering A 676 (2016) 377–384
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Low-cycle fatigue behavior of a cast Mg–Y–Nd–Zr alloy by T6 heat treatment Huizhong Li a,b, Feng Lv a, Zeyi Xiao c, Xiaopeng Liang a,b,n, Fengjian Sang a, Pengwei Li a a
School of Materials Science and Engineering, Central South University, Changsha 410083, China State Key Laboratory of Powder Metallurgy, Central South University, Changsha 410083, China c School of Metallurgy and Environment, Central South University, Changsha 410083, China b
art ic l e i nf o
a b s t r a c t
Article history: Received 5 July 2016 Received in revised form 31 August 2016 Accepted 1 September 2016 Available online 2 September 2016
Strain controlled low-cycle fatigue (LCF) tests of a cast Mg-3.5Y-2.4Nd-0.5Zr alloy by two different T6 heat treatment (T60: solution at 525 °C for 4 h and then aging at 225 °C for 34 h; T61: conditional industrial T6 heat treatment - solution at 520 °C for 8 h and aging at 225 °C for 16 h) were studied at room temperature under different total strain amplitudes. And theirs fatigue parameters, fatigue lives and failure behaviors after cyclic loading were evaluated. At relatively low total strain amplitudes (0.3–0.45%), the T60 alloy presented basically cyclic stabilization; and the T61 alloy was increasing hardening until failure. While at higher total strain amplitudes (0.45–0.9%), both them showed hardening through the entire LCF testing. Fatigue cracks of the alloy initiated from the specimen surface, the fracture surfaces became more rough and concavo-convex as the strain amplitude increased. Due to the fine and dense precipitates, the T60 alloy showed the super fatigue properties. And the alloy after T60 heat treatment obtained a distinctly longer fatigue life than that of the alloy treated under T61 condition at any given total strain amplitude. & 2016 Elsevier B.V. All rights reserved.
Keywords: Magnesium alloy Low-cycle fatigue Heat treatment Fracture Precipitate
1. Introduction Magnesium (Mg) alloys, as the lightest weight structural metal material; they also have other excellent properties, such as high specific strength and stiffness, machinability and recyclability. These advantages have made Mg alloys more attractive than conventional structural materials (steels and aluminum alloys) in the transportation industry, in particular, as basic components in aircrafts or ground vehicles for which weight saving is vitally important [1]. However, these components are unavoidably used under service condition of cyclic deformation or vibration [2], leading to catastrophic fracture after certain period of time. Accordingly, the fatigue behavior of these materials needs to be further investigated for safety and reliability requirements. The fatigue properties of Mg alloys produced by various techniques, such as rolling, casting and extrusion, have been studied in the past. For example, Luke H. Rettberg et al. [3] studied the LCF fatigue behavior of AZ91 and AM60 magnesium alloys, which was processed by the super vacuum die-casting method. They reported that after T6 heat treatment, the fatigue cracks initiated from n Corresponding author at: School of Materials Science and Engineering, Central South University, Changsha 410083, China. E-mail addresses: [email protected], [email protected] (X. Liang).
http://dx.doi.org/10.1016/j.msea.2016.09.001 0921-5093/& 2016 Elsevier B.V. All rights reserved.
either the surface or subsurface pores in all tested specimens. Z.Y. Nan and S. Ishihara [4] investigated crack initiation and propagation behaviors during fatigue of the extruded magnesium alloy AZ31 in detail and discussed their relationship with fatigue lives. In the research into the LCF of an extruded Mg-3Nd-0.2Zn-0.5Zr magnesium alloy [5], results showed that the alloy exhibited basically asymmetrical hysteresis loops; and the cyclic deformation characteristics showed cyclic softening at low strain amplitudes while cyclic hardening at high strain amplitudes. Sung Hyuk Park et al. [6] discussed the LCF characteristics of rolled Mg–3Al–1Zn alloy along the rolling direction, and they found that the alloy had a strong basal texture, therefore, the fatigue deformation was predominated by the alternation of twinning and detwinning during each cycle, which made the cyclic stress response unstable. With solid solution and precipitation hardening, and particularly the precipitation of a fine dispersion of inter-metallic particles, Mg alloys containing rare earth (RE) elements have the best combination of mechanical properties and corrosion resistance [7– 9]. As is reported in our earlier study [10], microstructure and mechanical properties of a Mg-Y-Nd-Zr alloy have been investigated. And we did find that after solution treatment at 525 °C for 4 h and subsequent aging at 225 °C for 34 h, the alloy displayed remarkable mechanical properties with a UTS of 301 MPa. This result was 20% higher compared with the alloy treated under T61 condition (solution treatment at 520 °C for 8 h and subsequent
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aging at 225 °C for 16 h). In this paper, we will study the fatigue resistance to find out whether it was also improved. The main aim of the present investigation are to confirm the cyclic deformation characteristics (cyclic hardening or softening) of the Mg-3.5Y-2.4Nd-0.5Zr alloy, evaluate the fatigue life under varying strain amplitudes, and obtain cyclic parameters. The failure behaviors after cyclic loading of tested alloy will be highly concerned in order to gain deep insights into their fatigue failure mechanisms. The effect of the two heat treatment (T60 and T61) on the fatigue life of the alloy will be studied; and the reasons for longer fatigue life of T60 alloy will also be discussed. The results of this work might provide a broader perspective for the development of the Mg-Y-Nd-Zr alloy.
2. Material and experimental procedures The material investigated was a low-pressure casting Mg alloy, with nominal composition: Mg, 3.5 wt% Y, 2.4 wt% Nd, and 0.5 wt% Zr. Specimens were cut from the ingot by electric spark linear cutting machine. The as-cast samples were heat-treated using T6 technology (T60: solution treatment for 4 h at 525 °C and then aging treatment for 34 h at 225 °C; T61: solution treatment for 8 h at 520 °C and then aging treatment for 16 h at 225 °C). The temperatures were controlled within 71 °C of the desired temperatures during the heating process. Tensile tests were performed at a rate of 2 mm/min on a MTS 810 machine at ambient temperature. The metallographic microstructures of the specimens prepared by standard techniques and etched in 4% HNO3 with ethanol were observed by optical microscope (OM, XJP-6A). The characterization of phases after heat treatment was carried out on transmission electron microscopy (TEM, Tencnai G220) operated at 200 kV. TEM foils were prepared after mechanical grinding to 0.08 mm, followed by twin-jet electrical polishing in a solution of 4% perchloric acid and 96% ethanol at 35 V and 20 °C. The fracture surfaces after fatigue tests were examined with a scanning electron microscope (SEM, FEI Quanta-200) to identify fatigue crack initiation sites and propagation characteristics. Strain-controlled, pull-push type fatigue tests were performed using a MTS 810 fatigue testing machine with MTS Test Suite control software. Sinusoidal waveform loading with a zero mean strain (i.e., a strain ratio of R ε= − 1, completely reversed strain cycle) and a constant frequency of 0.5 Hz was applied during the cyclic deformation tests. The strain was also measured by a clip-on 25 mm extensometer attached to the gage length. Low cycle fatigue tests were performed at room temperature of 25 °C for different strain amplitudes of 0.3%, 0.35%, 0.45%, 0.6%, 0.75%, 0.8% and 0.9%. At least two samples were tested at each set of testing parameters. Failure was defined as complete specimen separation or 20% tensile load drop, whichever occurred first. Before testing, the gage surface of all fatigue samples were polished with 1600 grit silicon carbide abrasive paper to avoid the influence of machining on the fatigue result.
3. Results 3.1. Tensile properties The tensile true stress-true strain curves of the alloys at room temperature are shown in Fig. 1 and the tensile properties obtained are listed in the table. As can be seen in the table, the T61 alloy displays a yield strength (YS) of 166 MPa, an ultimate tensile strength (UTS) of 273 MPa and an elongation of 4.1%. In contrast with the T61 alloy, obvious improvements of the YS and UTS are
Fig. 1. The tensile true stress-strain curves of the T60 alloy and T61 alloy tested at a rate of 2 mm/min.
observed in T60 alloy; and the T60 alloy shows significantly higher YS of 200 MPa, UTS of 301 MPa, and an elongation of 6.9%. 3.2. Fatigue properties 3.2.1. Stress response during cyclic deformation For the studied alloy in the T60 and T61 heat treated condition, the evolution of stress amplitudes with respect to the number of cycles to failure at different total strain amplitudes on a semi-log scale during LCF tests are shown in Fig. 2. In general, with increasing total strain amplitudes, the stress amplitudes increased and the fatigue life of the alloy decreased. Depending on the test condition and initial state, a metal could go through cyclic hardening, cyclic softening, or keep cyclically stable. Under a total strain control condition, cyclic hardening would lead to an increasing peak stress (or stress amplitude) with increasing number of cycles. While cyclic softening would be characterized by decreasing stress amplitude and increasing plastic strain amplitude, which would lead to early failure [11]. In the study of T60 alloy, the stress amplitudes remains almost constant during cyclic deformation at the low total strain amplitude of 0.3%. With increasing total strain amplitudes, cyclic hardening occurs. As the applied total strain amplitude increases from 0.35% up to 0.45%, the alloy shows basically cyclic stabilization within the initial few cycles, followed by a little cyclic hardening until failure. As indicated by the increasing slope in the semi-log scale diagram (Fig. 2(a)), the tendency toward cyclic hardening becomes stronger with increasing the total strain amplitude from 0.6% to 0.75%. At the total strain amplitudes between 0.6% and 0.9%, the alloy is increasing hardening until failure. Another notable change should be recorded on the plot is that the threshold cyclic stress amplitudes increases from 122 to 213 MPa with increasing total strain amplitudes from 0.3% to 0.9%. As reported by S.K. Shaha [12], the cyclic hardening behavior of an alloy mainly depends on the yield strength of the material. It means that if the total stress amplitude of the material is much lower than the yield strength, the alloy will not undergo a significant cyclic hardening and plastic deformation during cyclic deformation. Corresponding to Fig. 2(a), at lower total strain amplitudes 0.3–0.35%), the maximum total stress amplitude is much lower (122–128 MPa) than the yield strength (200 MPa), so the studied alloy does not show hardening apparently. While at higher total strain amplitudes 0.45–0.9%), the material shows hardening throughout its fatigue life, with the maximum total stress amplitude much closer (152 MPa–213 MPa) to the yield strength (200 MPa) of the alloy. However, at T61 condition (Fig. 2(b)), the cyclic stress response behavior is similar to T60 treatment at higher total strain
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Fig. 2. Stress amplitude versus the number of cycles of the (a) T60 and (b) T61 heat treated alloy tested at different total strain amplitudes at a strain ratio of R ε= − 1.
amplitudes (0.45–0.9%), while at lower total strain amplitudes (0.3–0.35%), continuous cyclic hardening is also seen through the entire LCF testing, which is different from the T60 condition. And the threshold cyclic stress amplitudes of T61 alloy increases from 109 to 199 MPa with increasing total strain amplitudes from 0.3% to 0.9%, respectively. 3.2.2. Hysteresis loops Fig. 3(a) shows the first cycle and mid-life cycle hysteresis loops at a total strain amplitude of 0.75%. It is seen that for this T60 treated Mg alloy, the hysteresis loops are basically symmetrical at the beginning of the fatigue life and at the mid-life. This suggests that the tensile and compressive deformation is nearly the same during cyclic straining, which is in sharp contrast to the asymmetric hysteresis loops that were presented in the extruded Mg alloys [13–16]. Compared with the first loop, the mid-life loop at a total strain amplitude of 0.75% shows a higher peak stress reflecting the occurrence of cyclic hardening, which corresponds well to the stress response shown in Fig. 2(a). And Fig. 3(b) shows the typical stress-strain hysteresis loops of the T60 alloy for the mid-life cycle at different applied total strain amplitudes in the LCF tests. Obviously, with the increasing total strain amplitude, the mid-life minimum and maximum peak stress increases and so does the width and height of the hysteresis loops. Thus, the hysteresis loops exhibited a characteristic clockwise rotation. 3.2.3. Fatigue Life and LCF parameters The cyclic strain-life ( ∆ε−Nf ) correlation is an important factor to evaluate the fatigue properties of materials, and is also a crucial aspect to improve the fatigue service performance. Fig. 4 shows
Fig. 4. Fatigue life of the alloy in the T60 heat treated condition in comparison with the same alloy treated under T61 condition.
the fatigue lifetime curves for the alloy in the present T60 tempered condition along with the same alloy treated under T61 condition. It can be found that under different heat treatment conditions, the fatigue lives of this alloy all increases with decreasing total strain amplitudes. And at all levels of total strain amplitude, the fatigue life of the T60 alloy appears significantly longer than that of the alloy under T61 condition. The total strain amplitude ( ∆ε/2) could be expressed as elastic strain amplitude (∆εe /2) and plastic strain amplitude (∆εp/2) [17];
Fig. 3. Typical hysteresis loops of the T60 alloy of (a) the first and mid-life cycles at a total strain amplitude of 0.75%, and (b) the mid-life cycles at different total strain amplitudes. Notice the clockwise rotation as the strain amplitude increases and the nonlinear behavior after load reversal.
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( ∆ε/2) = ( ∆εe/2)+(∆εp /2)
(1)
The first part could further be represented by the well-known Basquin equation as follow,
∆εe σ′(2Nf )b = f 2 E
(2)
where E is the Young's modulus (for the WE43 the value is approximately 40 GPa), Nf is the fatigue life or number of cycles to failure, σ′f is the coefficient of fatigue strength, b is the exponent of fatigue strength of alloy. And the second term of Eq. (1) could be replaced by the Coffin– Manson relation,
∆εp 2
c = ε′( f 2Nf )
(3)
where ε′f denote the fatigue plasticity coefficient, and c is the fatigue plasticity exponent. For most metals, b≈−0. 07~ − 0. 15, c≈−0. 5~ − 0. 7, high b value means the high strength and high c value indicates the good plasticity. On the other hand, the cyclic stress-strain response is also an important material property in designing for enhanced fatigue resistance. It described the relationship between flow stress and plastic strain amplitude under cyclic loading. The cyclic stressstrain curve might be described by the following equation:
(
∆σ /2 = K′ ∆εp /2
n′
)
(4)
where K′ is the cyclic strength coefficient and n′ is the cyclic strainhardening exponent. The stress amplitude and plastic strain amplitude in Eq. (4) corresponds to those at the half life. The variation of the number of reversals to failure (2Nf) with the total strain amplitude, together with elastic and plastic strain amplitude at the mid-life, are shown in Fig. 5. As is depicted in this figure, when the (elastic, plastic, and total) strain amplitude is lower, the fatigue lifetime appears higher; the experimental data obtained over the wide range of strain amplitudes followed both the Basquin and Coffin–Manson relationships well. And the relationship between cyclic stress amplitude and plastic strain amplitude of the alloy are shown in Fig. 6. In this study, the value of elastic or plastic strain amplitudes are taken at mid-life from the hysteresis loops for the tested alloy considering the stable condition of the LCF. According to Eq. (2)– (4), the plots of Figs. 5 and 6 are used to calculate the LCF parameters. The evaluated LCF parameters are determined using linear regression analysis and summarized in Table 1. It can be noticed
that the T60 alloy always exhibits higher LCF parameters compared to the T61 alloy. Specifically, the fatigue strength coefficient σ′f , and the fatigue strength exponent b significantly increases to 520 MPa and 0.171, respectively. In comparison with the T61 alloy, the fatigue ductility coefficient and exponent of the T60 alloy both increases from 2.7% and 0.485 to 9.1% and 0.675. At the same time, the cyclic strain hardening exponent, n′ and the cyclic strength coefficient, K′ for the T60 alloy are 0.167 and 569.737 MPa, respectively, which are also a little higher than those for the T61 alloy. Thus, it could be concluded that the T60 alloy has significantly longer fatigue life than that of the T61 alloy. 3.3. Fractography Generally, the LCF life of a fatigue failure sample consists of crack-nucleating and crack-propagating life. The nucleation of the fatigue micro-crack initiator presents some primary patterns, such as the cleavage of surface slip-bands, the breakaway of inclusions apart from the interface of based phase, as well as the cleavage of grain boundaries. For the smooth LCF samples of the die-cast Mg alloys, the micro-crack mainly initiates at the machining defects of samples surface such as the scratches and corners, or the casting defects inside samples such as the pores, the inclusions or the unmelted hardening phases. Because those locations favorably give rise to the stress and strain concentration of local micro-regions, the micro-crack easily nucleates and continuously grows and propagates, finally leading to its breakaway from the based phase and the fracture of materials. Fig. 7 shows the typical SEM images of the fracture surfaces of the T60 specimens tested at a total strain amplitude of 0.3%, 0.6% and 0.9%, respectively. And can be concluded that regardless of the total strain amplitudes, the fatigue crack initiated from the specimen surface or near-surface defects. At low strain amplitude of 0.3%, there appear some flats (Fig. 7(a)) on the fracture surface. However, at a relatively high strain amplitude of 0.6%, the alloy has a rougher fracture surface (Fig. 7(c)). As is indicated by the dashed line on the images, with decreasing total strain amplitudes, the area of the fatigue crack propagation zone increased. This corresponded to the extent of the increase process in the plastic strain amplitude prior to failure [13]. As a result, the decreased area of crack propagation zone leads to shortened fatigue lifetime. As is obviously observed from Fig. 7(b) and (d), there exist some fatigue striation-like features on the flats in the fatigue crack propagation zone, and each striation exhibits a propagation period of fatigue crack under cyclic loading. The appearance of striation means the increase of the fatigue crack propagation resistance,
Fig. 5. Total, elastic, and plastic strain amplitudes vs. the number of reversals to failure from the alloy: (a) is the alloy with T60 treated; and (b) is the alloy with T61 treated. Log-log linear regression fits were used to determine the fatigue parameters.
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Fig. 6. Cyclic stress amplitude vs. plastic strain amplitude of the alloy tested at room temperature: (a) is the alloy with T60 treated; and (b) is the alloy with T61 treated.
Table 1 Low-cycle fatigue parameters obtained for the alloy. Low-cycle fatigue parameters
T60 Mg alloy
T61 Mg alloy
Cyclic strain hardening exponent, n′ Cyclic strength coefficient, K′, MPa Fatigue strength coefficient, σ′f , MPa Fatigue strength exponent, b Fatigue ductility coefficient, ε′f Fatigue ductility exponent, c
0.167 569.7 520 0.171 0.091 0.675
0.157 516.8 440 0.162 0.027 0.485
and thus results in a longer fatigue life. And the spacing of fatigue striations increases with the total strain amplitude increasing. Therefore, the larger spacing of fatigue striations at the higher total strain amplitudes showing a shorter fatigue life, which corresponds to the fast crack propagation. At low strain amplitude, the micro-crack usually nucleate from the defects such as inclusions, pores, notches, as well as nonuniform deformation of grains, and the failure exhibits a typical multi-stage fatigue cracks growth feature [18–20]. With the number of cycles increasing, once the micro-cracks initiate, the multi-stage fatigue cracks will propagate and grow through the Mg(α) matrix. For each stage of the fatigue crack, the dislocation will occur to cross slip, pile up and entangle inside the crack-tips plastic region. As a consequence, once fracture takes place in the alloys after bearing a high number of cycles, the fatigue flats and striations, even the fatigue slip-bands, can be distinctly observed on the fracture surfaces. As is shown at high strain amplitude (Fig. 7(f)), the surface are composed of tear ridges and cleavage planes, suggesting its similarity to the tensile fracture. At high strain amplitude, the microcrack primarily initiates from the grain boundaries and propagates along the grain boundaries and the eutectic phase boundaries, where the incompatibility of local deformation is rather evident and the crack propagation resistance is also low. Once the microcrack initiates, it will rapidly propagate, displaying the mono-crack initiator characteristics and resulting in a shorter LCF life. So the fatigue fracture is similar to the tensile fracture and the fatigue striations are difficult to be detected.
condition. In general, the fatigue properties of the alloy are mainly determined by the following aspects: (1) Sample size: fatigue strength decreases with the increase of sample size. (2) Microstructures: the reduction of grain size could introduce more grain boundaries, which will impede the dislocation movement in the process of deformation, and further results the enhanced resistance of crack initiation. Thus, the fatigue properties improve as grain size reduces. (3) Surface states: with rough surfaces, the sample under loading will lead to high stress concentration at there and simulate to the nucleation of micro-cracks. (4) Environmental factors: such as temperature, corrosion environment, loading ways and so on. And this paper mainly analyze the influence from the microstructures. The optical microstructures of the alloy under T60 and T61 condition are shown in Fig. 8. It can be found that the average grain size of the alloy after T60 heat treatment is about 55 μm ; while the grain size of T61 alloy is about 60 μm . This phenomenon maybe attributed to the longer solid-solution time in the T61 alloy. And this subtle distinctions will not have a great impact on the fatigue properties. The Mg-Y-Nd-Zr alloy is a precipitation hardened alloy, and the precipitation sequence consists of the following steps: super00 0 saturated α-Mg solid solution-β (DO19)-β (bcc)-β1 (fcc) - β (fcc) [21–25]. With different heat treatment condition, the number density, distribution, kinds and size of the precipitations are also diverse, and further lead to the different mechanical properties. The dark field (DF) TEM micrographs and corresponding selected area election diffraction (SAED) patterns of the T60 alloy and the T61 alloy are shown in Fig. 9. According to Fig. 9(a), the lamellar precipitates have an average size about 40 nm in length, which ̅ ⎤⎦ direction of redistributed densely and uniformly along ⎡⎣ 1100 ciprocal space. In the corresponding SAED pattern (Fig. 9(b)), there ̅ ⎤⎦ . Such a typical DO19 structure are diffraction spots at 1/2 ⎡⎣ 1100 00
β″
According to Fig. 4, the alloy under T60 condition shows an obvious longer fatigue life than that of the alloy under T61
α
̅ ) //( 1100 ̅ ) . Furthermore, there exist reflections at ( 1100 β″ α ̅ } , 2/4{ 1100 ̅ } and 3/4{ 1100 ̅ } and the phase seems to 1/4{ 1100 α α α be
β′ phase, with lattice parameters of a = 2×aα− Mg=0. 6418nm,
(
4. Discussion
α
can be identified to be β phase, with lattice parameters a = 2×aα− Mg=0. 6418nm , c=cα− Mg=0. 521nm [26] and an orienta⎡⎣ 1120 ̅ ⎤⎦ //⎡⎣ 1120 ̅ ⎤⎦ , tion relationship with the matrix is
)
̅ =2. 223nm , c=cα− Mg=0. 521nm [27]. Therefore, it b = 8 × d 1010 0 00 indicates that β phase coexists with β phase after T60 heat treatment. When the alloy is treated under T61 condition (Fig. 9 (c) and (d)), the kinds of precipitations are the same with T60 alloy, but distributions of these precipitates at the micro scale are inhomogeneous. All these maybe attributed to the longer aging time
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Fig. 7. SEM micrographs of the fracture surfaces of the T60 alloy fatigued at different total strain amplitudes: (a and b) 0.3%, (c and d) 0.6% and (e and f) 0.9%.
in the T60 alloy. The process of forming β phase is slow in a re00 latively low temperature. But as time goes on, most of β phase 0 will transformed into β phase. So the alloy under T60 condition 0 has a high volume fraction of β phase, which has an excellent resistance to the basal slip in the Mg alloy [28]. Therefore, its strengthening effect is developed. By comparison, the alloy under T60 condition can obtain finer and denser precipitates, which will act as the potential obstacles to dislocations. When the samples under loading, the dislocations will pile up in the precipitates, and then prompt to the nucleation of micro-cracks. In this case, the fracture will start from grain 0
interiors, and expand to the grain boundaries. Due to the hindering effect of the grain boundaries, the fracture will be delayed. As a result, the alloy shows a longer fatigue life. Whereas, with low precipitates volume fraction, the dislocations will bypass the precipitates and then pile up at the grain boundaries. Due to the high stress concentration at there, the fracture will be accelerated. With the number of fatigue cycles increasing, the dislocation density gradually increases. And the dislocation entanglement becomes more chaotic, which augments the resistance of subsequent dislocation movement. Moreover, the effect of the solidification microstructure on the cyclic hardening behavior is mainly
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Fig. 8. The optical microstructures of the alloy: (a) T60 and (b) T61.
attributed to the interaction between the dislocation and the precipitates or the grains boundary. When cyclic deformation takes place in the alloy under alternating loading, the cross-slip ability of dislocation will continuously increase [29]. The resistance of dislocation slipping is resulted from the boundaries between the precipitations and the grains, in addition, the refiner grains inevitably shorten the mean free path of dislocation slipping and thus lead to the higher cyclic hardening ability. The Mg-Y-Nd-Zr alloy has refine grains, better grain distributions and the better morphological aspects of phase, so the dislocations easily pile up, entangle and reach to a quasi-stable saturation state in the low strain range.
5. Conclusions In this paper, the low-cycle fatigue properties of the Mg-Y-NdZr alloy by axial cyclic loading under different total strain amplitudes are investigated. And the following conclusions can be drawn from this work: (1) Compared to the heat treatment under T61 condition, the Mg3.5Y-2.4Nd-0.5Zr alloy treated under T60 condition not only could enhance the strength, but also obtain a longer fatigue life.
̅ ⎤⎦ ) and T61 alloy ((c) TEM dark field and (d) SAED pattern with B// ⎡⎣ 1120 ̅ ⎤⎦ ). Fig. 9. TEM images for the T60 alloy ((a) TEM dark field and (b) SAED pattern with B// ⎡⎣ 1120
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(2) During low total strain amplitude (0.3–0.45%), the T60 alloy almost presented cyclic stabilization; and the T61 alloy was increasing hardening until failure. While at higher total strain amplitudes (0.45–0.9%), both them showed continuous hardening. (3) Fatigue cracks of the alloy initiated from the specimen surface, as the strain amplitude increased, the fracture surface became more rough and concavo-convex. (4) The alloy after T60 heat treatment achieved a substantially longer fatigue life than that of the alloy treated under T61 condition at any given total strain amplitude. The improved fatigue resistance of the T60 alloy was mainly attributed to the fine and dense precipitates.
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cast Mg-Y-Nd-Zr alloy, Mater. Sci. Eng. A 667 (2016) 409–416. [11] George Ellwood Dieter, Mechanical Metallurgy, 3rd ed, McGraw-Hill, Boston, USA, 1986. [12] S.K. Shaha, F. Czerwinski, W. Kasprzak, J. Friedman, D.L. Chen, Monotonic and cyclic deformation behavior of the Al–Si–Cu–Mg cast alloy with micro-additions of Ti, V and Zr, Int. J. Fatigue 70 (2015) 383–394. [13] X.Z. Lin, D.L. Chen, Strain controlled cyclic deformation behavior of an extruded magnesium alloy, Mater. Sci. Eng. A 496 (2008) 106–113. [14] S. Begum, D.L. Chen, S. Xu, A.A. Luo, Low cycle fatigue properties of an extruded AZ31 magnesium alloy, Int. J. Fatigue 31 (2009) 726–735. [15] S. Begum, D.L. Chen, S. Xu, A.A. Luo, Strain-controlled low-cycle fatigue properties of a newly developed extruded magnesium alloy, Met. Mater. Trans. A 39 (2008) 3014–3026. [16] S. Begum, D.L. Chen, S. Xu, A.A. Luo, Effect of strain ratio and strain rate on low cycle fatigue behavior of AZ31 wrought magnesium alloy, Mater. Sci. Eng. A 517 (2009) 334–343. [17] G.E. Dieter, Mechanical Metallurgy, McGraw-Hill Series, McGraw-Hill, New York, NY, 1986. [18] J. Schijve, Significant of fractography for investigations of fatigue crack growth under variable amplitude loading, Fatigue Fract. Eng. Mater. Struct. 22 (1999) 87–99. [19] Y. Xue, H. El Kadiri, M.F. Horstemeyer, Micromechanisms of multistage fatigue crack growth in a high-strength aluminum alloy, Acta Mater. 55 (2007) 1975–1984. [20] Y.B. Xue, Modeling fatigue small-crack growth with confidence-A multistage approach, Int. J. Fatigue 32 (2010) 1210–1219. [21] R.L. Xin, L. Li, K. Zeng, Bo Song, Q. Liu, Structural examination of aging precipitation in a Mg–Y–Nd alloy at different temperatures, Mater. Charact. 62 (2011) 535–539. [22] Y. Gao, H. Liu, R. Shi, N. Zhou, Z. Xu, Y.M. Zhu, J.F. Nie, Y. Wang, Simulation study of precipitation in an Mg–Y–Nd alloy, Acta Mater. 60 (2012) 4819–4832. [23] G. Barucca, R. Ferragut, F. Fiori, D. Lussana, P. Mengucci, F. Moia, G. Riontino, Formation and evolution of the hardening precipitates in a Mg-Y-Nd alloy, Acta Mater. 59 (2011) 4151–4158. [24] J.F. Nie, B.C. Muddle, Precipitation in magnesium alloy WE54 during isothermal ageing at 250 °C, Scr. Mater. 40 (1999) 1089–1094. [25] J.F. Nie, B.C. Muddle, Characterisation of strengthening precipitate phases in a Mg-Y-Nd alloy, Acta Mater. 48 (2000) 1691–1703. [26] B. Smola, I. Stulı ́ková, F. von Buch, B.L. Mordike, Structural aspects of high performance Mg alloys design, Mater. Sci. Eng. A 324 (2002) 113–117. [27] J.F. Nie, X.L. Xiao, C.P. Luo, B.C. Muddle, Characterisation of precipitate phases in magnesium alloys using electron microdiffraction, Micron 32 (2001) 857–863. [28] P.H. Fu, L.M. Peng, H.Y. Jiang, Z.Y. Zhang, C.Q. Zhai, Fracture behavior and mechanical properties of Mg–4Y–2Nd–1Gd–0.4Zr (wt%) alloy at room temperature, Mater. Sci. Eng. A 486 (2008) 572–579. [29] J.A. Robert, K. Yashashree, Are rate sensitivity and strength effected by crossslip in nano-twinned fcc metals, Scr. Mater. 58 (2008) 389–392.
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Materials Science & Engineering A journal homepage: www.elsevier.com/locate/msea
Low-cycle fatigue behavior of a cast Mg–Y–Nd–Zr alloy by T6 heat treatment Huizhong Li a,b, Feng Lv a, Zeyi Xiao c, Xiaopeng Liang a,b,n, Fengjian Sang a, Pengwei Li a a
School of Materials Science and Engineering, Central South University, Changsha 410083, China State Key Laboratory of Powder Metallurgy, Central South University, Changsha 410083, China c School of Metallurgy and Environment, Central South University, Changsha 410083, China b
art ic l e i nf o
a b s t r a c t
Article history: Received 5 July 2016 Received in revised form 31 August 2016 Accepted 1 September 2016 Available online 2 September 2016
Strain controlled low-cycle fatigue (LCF) tests of a cast Mg-3.5Y-2.4Nd-0.5Zr alloy by two different T6 heat treatment (T60: solution at 525 °C for 4 h and then aging at 225 °C for 34 h; T61: conditional industrial T6 heat treatment - solution at 520 °C for 8 h and aging at 225 °C for 16 h) were studied at room temperature under different total strain amplitudes. And theirs fatigue parameters, fatigue lives and failure behaviors after cyclic loading were evaluated. At relatively low total strain amplitudes (0.3–0.45%), the T60 alloy presented basically cyclic stabilization; and the T61 alloy was increasing hardening until failure. While at higher total strain amplitudes (0.45–0.9%), both them showed hardening through the entire LCF testing. Fatigue cracks of the alloy initiated from the specimen surface, the fracture surfaces became more rough and concavo-convex as the strain amplitude increased. Due to the fine and dense precipitates, the T60 alloy showed the super fatigue properties. And the alloy after T60 heat treatment obtained a distinctly longer fatigue life than that of the alloy treated under T61 condition at any given total strain amplitude. & 2016 Elsevier B.V. All rights reserved.
Keywords: Magnesium alloy Low-cycle fatigue Heat treatment Fracture Precipitate
1. Introduction Magnesium (Mg) alloys, as the lightest weight structural metal material; they also have other excellent properties, such as high specific strength and stiffness, machinability and recyclability. These advantages have made Mg alloys more attractive than conventional structural materials (steels and aluminum alloys) in the transportation industry, in particular, as basic components in aircrafts or ground vehicles for which weight saving is vitally important [1]. However, these components are unavoidably used under service condition of cyclic deformation or vibration [2], leading to catastrophic fracture after certain period of time. Accordingly, the fatigue behavior of these materials needs to be further investigated for safety and reliability requirements. The fatigue properties of Mg alloys produced by various techniques, such as rolling, casting and extrusion, have been studied in the past. For example, Luke H. Rettberg et al. [3] studied the LCF fatigue behavior of AZ91 and AM60 magnesium alloys, which was processed by the super vacuum die-casting method. They reported that after T6 heat treatment, the fatigue cracks initiated from n Corresponding author at: School of Materials Science and Engineering, Central South University, Changsha 410083, China. E-mail addresses: [email protected], [email protected] (X. Liang).
http://dx.doi.org/10.1016/j.msea.2016.09.001 0921-5093/& 2016 Elsevier B.V. All rights reserved.
either the surface or subsurface pores in all tested specimens. Z.Y. Nan and S. Ishihara [4] investigated crack initiation and propagation behaviors during fatigue of the extruded magnesium alloy AZ31 in detail and discussed their relationship with fatigue lives. In the research into the LCF of an extruded Mg-3Nd-0.2Zn-0.5Zr magnesium alloy [5], results showed that the alloy exhibited basically asymmetrical hysteresis loops; and the cyclic deformation characteristics showed cyclic softening at low strain amplitudes while cyclic hardening at high strain amplitudes. Sung Hyuk Park et al. [6] discussed the LCF characteristics of rolled Mg–3Al–1Zn alloy along the rolling direction, and they found that the alloy had a strong basal texture, therefore, the fatigue deformation was predominated by the alternation of twinning and detwinning during each cycle, which made the cyclic stress response unstable. With solid solution and precipitation hardening, and particularly the precipitation of a fine dispersion of inter-metallic particles, Mg alloys containing rare earth (RE) elements have the best combination of mechanical properties and corrosion resistance [7– 9]. As is reported in our earlier study [10], microstructure and mechanical properties of a Mg-Y-Nd-Zr alloy have been investigated. And we did find that after solution treatment at 525 °C for 4 h and subsequent aging at 225 °C for 34 h, the alloy displayed remarkable mechanical properties with a UTS of 301 MPa. This result was 20% higher compared with the alloy treated under T61 condition (solution treatment at 520 °C for 8 h and subsequent
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aging at 225 °C for 16 h). In this paper, we will study the fatigue resistance to find out whether it was also improved. The main aim of the present investigation are to confirm the cyclic deformation characteristics (cyclic hardening or softening) of the Mg-3.5Y-2.4Nd-0.5Zr alloy, evaluate the fatigue life under varying strain amplitudes, and obtain cyclic parameters. The failure behaviors after cyclic loading of tested alloy will be highly concerned in order to gain deep insights into their fatigue failure mechanisms. The effect of the two heat treatment (T60 and T61) on the fatigue life of the alloy will be studied; and the reasons for longer fatigue life of T60 alloy will also be discussed. The results of this work might provide a broader perspective for the development of the Mg-Y-Nd-Zr alloy.
2. Material and experimental procedures The material investigated was a low-pressure casting Mg alloy, with nominal composition: Mg, 3.5 wt% Y, 2.4 wt% Nd, and 0.5 wt% Zr. Specimens were cut from the ingot by electric spark linear cutting machine. The as-cast samples were heat-treated using T6 technology (T60: solution treatment for 4 h at 525 °C and then aging treatment for 34 h at 225 °C; T61: solution treatment for 8 h at 520 °C and then aging treatment for 16 h at 225 °C). The temperatures were controlled within 71 °C of the desired temperatures during the heating process. Tensile tests were performed at a rate of 2 mm/min on a MTS 810 machine at ambient temperature. The metallographic microstructures of the specimens prepared by standard techniques and etched in 4% HNO3 with ethanol were observed by optical microscope (OM, XJP-6A). The characterization of phases after heat treatment was carried out on transmission electron microscopy (TEM, Tencnai G220) operated at 200 kV. TEM foils were prepared after mechanical grinding to 0.08 mm, followed by twin-jet electrical polishing in a solution of 4% perchloric acid and 96% ethanol at 35 V and 20 °C. The fracture surfaces after fatigue tests were examined with a scanning electron microscope (SEM, FEI Quanta-200) to identify fatigue crack initiation sites and propagation characteristics. Strain-controlled, pull-push type fatigue tests were performed using a MTS 810 fatigue testing machine with MTS Test Suite control software. Sinusoidal waveform loading with a zero mean strain (i.e., a strain ratio of R ε= − 1, completely reversed strain cycle) and a constant frequency of 0.5 Hz was applied during the cyclic deformation tests. The strain was also measured by a clip-on 25 mm extensometer attached to the gage length. Low cycle fatigue tests were performed at room temperature of 25 °C for different strain amplitudes of 0.3%, 0.35%, 0.45%, 0.6%, 0.75%, 0.8% and 0.9%. At least two samples were tested at each set of testing parameters. Failure was defined as complete specimen separation or 20% tensile load drop, whichever occurred first. Before testing, the gage surface of all fatigue samples were polished with 1600 grit silicon carbide abrasive paper to avoid the influence of machining on the fatigue result.
3. Results 3.1. Tensile properties The tensile true stress-true strain curves of the alloys at room temperature are shown in Fig. 1 and the tensile properties obtained are listed in the table. As can be seen in the table, the T61 alloy displays a yield strength (YS) of 166 MPa, an ultimate tensile strength (UTS) of 273 MPa and an elongation of 4.1%. In contrast with the T61 alloy, obvious improvements of the YS and UTS are
Fig. 1. The tensile true stress-strain curves of the T60 alloy and T61 alloy tested at a rate of 2 mm/min.
observed in T60 alloy; and the T60 alloy shows significantly higher YS of 200 MPa, UTS of 301 MPa, and an elongation of 6.9%. 3.2. Fatigue properties 3.2.1. Stress response during cyclic deformation For the studied alloy in the T60 and T61 heat treated condition, the evolution of stress amplitudes with respect to the number of cycles to failure at different total strain amplitudes on a semi-log scale during LCF tests are shown in Fig. 2. In general, with increasing total strain amplitudes, the stress amplitudes increased and the fatigue life of the alloy decreased. Depending on the test condition and initial state, a metal could go through cyclic hardening, cyclic softening, or keep cyclically stable. Under a total strain control condition, cyclic hardening would lead to an increasing peak stress (or stress amplitude) with increasing number of cycles. While cyclic softening would be characterized by decreasing stress amplitude and increasing plastic strain amplitude, which would lead to early failure [11]. In the study of T60 alloy, the stress amplitudes remains almost constant during cyclic deformation at the low total strain amplitude of 0.3%. With increasing total strain amplitudes, cyclic hardening occurs. As the applied total strain amplitude increases from 0.35% up to 0.45%, the alloy shows basically cyclic stabilization within the initial few cycles, followed by a little cyclic hardening until failure. As indicated by the increasing slope in the semi-log scale diagram (Fig. 2(a)), the tendency toward cyclic hardening becomes stronger with increasing the total strain amplitude from 0.6% to 0.75%. At the total strain amplitudes between 0.6% and 0.9%, the alloy is increasing hardening until failure. Another notable change should be recorded on the plot is that the threshold cyclic stress amplitudes increases from 122 to 213 MPa with increasing total strain amplitudes from 0.3% to 0.9%. As reported by S.K. Shaha [12], the cyclic hardening behavior of an alloy mainly depends on the yield strength of the material. It means that if the total stress amplitude of the material is much lower than the yield strength, the alloy will not undergo a significant cyclic hardening and plastic deformation during cyclic deformation. Corresponding to Fig. 2(a), at lower total strain amplitudes 0.3–0.35%), the maximum total stress amplitude is much lower (122–128 MPa) than the yield strength (200 MPa), so the studied alloy does not show hardening apparently. While at higher total strain amplitudes 0.45–0.9%), the material shows hardening throughout its fatigue life, with the maximum total stress amplitude much closer (152 MPa–213 MPa) to the yield strength (200 MPa) of the alloy. However, at T61 condition (Fig. 2(b)), the cyclic stress response behavior is similar to T60 treatment at higher total strain
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Fig. 2. Stress amplitude versus the number of cycles of the (a) T60 and (b) T61 heat treated alloy tested at different total strain amplitudes at a strain ratio of R ε= − 1.
amplitudes (0.45–0.9%), while at lower total strain amplitudes (0.3–0.35%), continuous cyclic hardening is also seen through the entire LCF testing, which is different from the T60 condition. And the threshold cyclic stress amplitudes of T61 alloy increases from 109 to 199 MPa with increasing total strain amplitudes from 0.3% to 0.9%, respectively. 3.2.2. Hysteresis loops Fig. 3(a) shows the first cycle and mid-life cycle hysteresis loops at a total strain amplitude of 0.75%. It is seen that for this T60 treated Mg alloy, the hysteresis loops are basically symmetrical at the beginning of the fatigue life and at the mid-life. This suggests that the tensile and compressive deformation is nearly the same during cyclic straining, which is in sharp contrast to the asymmetric hysteresis loops that were presented in the extruded Mg alloys [13–16]. Compared with the first loop, the mid-life loop at a total strain amplitude of 0.75% shows a higher peak stress reflecting the occurrence of cyclic hardening, which corresponds well to the stress response shown in Fig. 2(a). And Fig. 3(b) shows the typical stress-strain hysteresis loops of the T60 alloy for the mid-life cycle at different applied total strain amplitudes in the LCF tests. Obviously, with the increasing total strain amplitude, the mid-life minimum and maximum peak stress increases and so does the width and height of the hysteresis loops. Thus, the hysteresis loops exhibited a characteristic clockwise rotation. 3.2.3. Fatigue Life and LCF parameters The cyclic strain-life ( ∆ε−Nf ) correlation is an important factor to evaluate the fatigue properties of materials, and is also a crucial aspect to improve the fatigue service performance. Fig. 4 shows
Fig. 4. Fatigue life of the alloy in the T60 heat treated condition in comparison with the same alloy treated under T61 condition.
the fatigue lifetime curves for the alloy in the present T60 tempered condition along with the same alloy treated under T61 condition. It can be found that under different heat treatment conditions, the fatigue lives of this alloy all increases with decreasing total strain amplitudes. And at all levels of total strain amplitude, the fatigue life of the T60 alloy appears significantly longer than that of the alloy under T61 condition. The total strain amplitude ( ∆ε/2) could be expressed as elastic strain amplitude (∆εe /2) and plastic strain amplitude (∆εp/2) [17];
Fig. 3. Typical hysteresis loops of the T60 alloy of (a) the first and mid-life cycles at a total strain amplitude of 0.75%, and (b) the mid-life cycles at different total strain amplitudes. Notice the clockwise rotation as the strain amplitude increases and the nonlinear behavior after load reversal.
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( ∆ε/2) = ( ∆εe/2)+(∆εp /2)
(1)
The first part could further be represented by the well-known Basquin equation as follow,
∆εe σ′(2Nf )b = f 2 E
(2)
where E is the Young's modulus (for the WE43 the value is approximately 40 GPa), Nf is the fatigue life or number of cycles to failure, σ′f is the coefficient of fatigue strength, b is the exponent of fatigue strength of alloy. And the second term of Eq. (1) could be replaced by the Coffin– Manson relation,
∆εp 2
c = ε′( f 2Nf )
(3)
where ε′f denote the fatigue plasticity coefficient, and c is the fatigue plasticity exponent. For most metals, b≈−0. 07~ − 0. 15, c≈−0. 5~ − 0. 7, high b value means the high strength and high c value indicates the good plasticity. On the other hand, the cyclic stress-strain response is also an important material property in designing for enhanced fatigue resistance. It described the relationship between flow stress and plastic strain amplitude under cyclic loading. The cyclic stressstrain curve might be described by the following equation:
(
∆σ /2 = K′ ∆εp /2
n′
)
(4)
where K′ is the cyclic strength coefficient and n′ is the cyclic strainhardening exponent. The stress amplitude and plastic strain amplitude in Eq. (4) corresponds to those at the half life. The variation of the number of reversals to failure (2Nf) with the total strain amplitude, together with elastic and plastic strain amplitude at the mid-life, are shown in Fig. 5. As is depicted in this figure, when the (elastic, plastic, and total) strain amplitude is lower, the fatigue lifetime appears higher; the experimental data obtained over the wide range of strain amplitudes followed both the Basquin and Coffin–Manson relationships well. And the relationship between cyclic stress amplitude and plastic strain amplitude of the alloy are shown in Fig. 6. In this study, the value of elastic or plastic strain amplitudes are taken at mid-life from the hysteresis loops for the tested alloy considering the stable condition of the LCF. According to Eq. (2)– (4), the plots of Figs. 5 and 6 are used to calculate the LCF parameters. The evaluated LCF parameters are determined using linear regression analysis and summarized in Table 1. It can be noticed
that the T60 alloy always exhibits higher LCF parameters compared to the T61 alloy. Specifically, the fatigue strength coefficient σ′f , and the fatigue strength exponent b significantly increases to 520 MPa and 0.171, respectively. In comparison with the T61 alloy, the fatigue ductility coefficient and exponent of the T60 alloy both increases from 2.7% and 0.485 to 9.1% and 0.675. At the same time, the cyclic strain hardening exponent, n′ and the cyclic strength coefficient, K′ for the T60 alloy are 0.167 and 569.737 MPa, respectively, which are also a little higher than those for the T61 alloy. Thus, it could be concluded that the T60 alloy has significantly longer fatigue life than that of the T61 alloy. 3.3. Fractography Generally, the LCF life of a fatigue failure sample consists of crack-nucleating and crack-propagating life. The nucleation of the fatigue micro-crack initiator presents some primary patterns, such as the cleavage of surface slip-bands, the breakaway of inclusions apart from the interface of based phase, as well as the cleavage of grain boundaries. For the smooth LCF samples of the die-cast Mg alloys, the micro-crack mainly initiates at the machining defects of samples surface such as the scratches and corners, or the casting defects inside samples such as the pores, the inclusions or the unmelted hardening phases. Because those locations favorably give rise to the stress and strain concentration of local micro-regions, the micro-crack easily nucleates and continuously grows and propagates, finally leading to its breakaway from the based phase and the fracture of materials. Fig. 7 shows the typical SEM images of the fracture surfaces of the T60 specimens tested at a total strain amplitude of 0.3%, 0.6% and 0.9%, respectively. And can be concluded that regardless of the total strain amplitudes, the fatigue crack initiated from the specimen surface or near-surface defects. At low strain amplitude of 0.3%, there appear some flats (Fig. 7(a)) on the fracture surface. However, at a relatively high strain amplitude of 0.6%, the alloy has a rougher fracture surface (Fig. 7(c)). As is indicated by the dashed line on the images, with decreasing total strain amplitudes, the area of the fatigue crack propagation zone increased. This corresponded to the extent of the increase process in the plastic strain amplitude prior to failure [13]. As a result, the decreased area of crack propagation zone leads to shortened fatigue lifetime. As is obviously observed from Fig. 7(b) and (d), there exist some fatigue striation-like features on the flats in the fatigue crack propagation zone, and each striation exhibits a propagation period of fatigue crack under cyclic loading. The appearance of striation means the increase of the fatigue crack propagation resistance,
Fig. 5. Total, elastic, and plastic strain amplitudes vs. the number of reversals to failure from the alloy: (a) is the alloy with T60 treated; and (b) is the alloy with T61 treated. Log-log linear regression fits were used to determine the fatigue parameters.
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Fig. 6. Cyclic stress amplitude vs. plastic strain amplitude of the alloy tested at room temperature: (a) is the alloy with T60 treated; and (b) is the alloy with T61 treated.
Table 1 Low-cycle fatigue parameters obtained for the alloy. Low-cycle fatigue parameters
T60 Mg alloy
T61 Mg alloy
Cyclic strain hardening exponent, n′ Cyclic strength coefficient, K′, MPa Fatigue strength coefficient, σ′f , MPa Fatigue strength exponent, b Fatigue ductility coefficient, ε′f Fatigue ductility exponent, c
0.167 569.7 520 0.171 0.091 0.675
0.157 516.8 440 0.162 0.027 0.485
and thus results in a longer fatigue life. And the spacing of fatigue striations increases with the total strain amplitude increasing. Therefore, the larger spacing of fatigue striations at the higher total strain amplitudes showing a shorter fatigue life, which corresponds to the fast crack propagation. At low strain amplitude, the micro-crack usually nucleate from the defects such as inclusions, pores, notches, as well as nonuniform deformation of grains, and the failure exhibits a typical multi-stage fatigue cracks growth feature [18–20]. With the number of cycles increasing, once the micro-cracks initiate, the multi-stage fatigue cracks will propagate and grow through the Mg(α) matrix. For each stage of the fatigue crack, the dislocation will occur to cross slip, pile up and entangle inside the crack-tips plastic region. As a consequence, once fracture takes place in the alloys after bearing a high number of cycles, the fatigue flats and striations, even the fatigue slip-bands, can be distinctly observed on the fracture surfaces. As is shown at high strain amplitude (Fig. 7(f)), the surface are composed of tear ridges and cleavage planes, suggesting its similarity to the tensile fracture. At high strain amplitude, the microcrack primarily initiates from the grain boundaries and propagates along the grain boundaries and the eutectic phase boundaries, where the incompatibility of local deformation is rather evident and the crack propagation resistance is also low. Once the microcrack initiates, it will rapidly propagate, displaying the mono-crack initiator characteristics and resulting in a shorter LCF life. So the fatigue fracture is similar to the tensile fracture and the fatigue striations are difficult to be detected.
condition. In general, the fatigue properties of the alloy are mainly determined by the following aspects: (1) Sample size: fatigue strength decreases with the increase of sample size. (2) Microstructures: the reduction of grain size could introduce more grain boundaries, which will impede the dislocation movement in the process of deformation, and further results the enhanced resistance of crack initiation. Thus, the fatigue properties improve as grain size reduces. (3) Surface states: with rough surfaces, the sample under loading will lead to high stress concentration at there and simulate to the nucleation of micro-cracks. (4) Environmental factors: such as temperature, corrosion environment, loading ways and so on. And this paper mainly analyze the influence from the microstructures. The optical microstructures of the alloy under T60 and T61 condition are shown in Fig. 8. It can be found that the average grain size of the alloy after T60 heat treatment is about 55 μm ; while the grain size of T61 alloy is about 60 μm . This phenomenon maybe attributed to the longer solid-solution time in the T61 alloy. And this subtle distinctions will not have a great impact on the fatigue properties. The Mg-Y-Nd-Zr alloy is a precipitation hardened alloy, and the precipitation sequence consists of the following steps: super00 0 saturated α-Mg solid solution-β (DO19)-β (bcc)-β1 (fcc) - β (fcc) [21–25]. With different heat treatment condition, the number density, distribution, kinds and size of the precipitations are also diverse, and further lead to the different mechanical properties. The dark field (DF) TEM micrographs and corresponding selected area election diffraction (SAED) patterns of the T60 alloy and the T61 alloy are shown in Fig. 9. According to Fig. 9(a), the lamellar precipitates have an average size about 40 nm in length, which ̅ ⎤⎦ direction of redistributed densely and uniformly along ⎡⎣ 1100 ciprocal space. In the corresponding SAED pattern (Fig. 9(b)), there ̅ ⎤⎦ . Such a typical DO19 structure are diffraction spots at 1/2 ⎡⎣ 1100 00
β″
According to Fig. 4, the alloy under T60 condition shows an obvious longer fatigue life than that of the alloy under T61
α
̅ ) //( 1100 ̅ ) . Furthermore, there exist reflections at ( 1100 β″ α ̅ } , 2/4{ 1100 ̅ } and 3/4{ 1100 ̅ } and the phase seems to 1/4{ 1100 α α α be
β′ phase, with lattice parameters of a = 2×aα− Mg=0. 6418nm,
(
4. Discussion
α
can be identified to be β phase, with lattice parameters a = 2×aα− Mg=0. 6418nm , c=cα− Mg=0. 521nm [26] and an orienta⎡⎣ 1120 ̅ ⎤⎦ //⎡⎣ 1120 ̅ ⎤⎦ , tion relationship with the matrix is
)
̅ =2. 223nm , c=cα− Mg=0. 521nm [27]. Therefore, it b = 8 × d 1010 0 00 indicates that β phase coexists with β phase after T60 heat treatment. When the alloy is treated under T61 condition (Fig. 9 (c) and (d)), the kinds of precipitations are the same with T60 alloy, but distributions of these precipitates at the micro scale are inhomogeneous. All these maybe attributed to the longer aging time
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Fig. 7. SEM micrographs of the fracture surfaces of the T60 alloy fatigued at different total strain amplitudes: (a and b) 0.3%, (c and d) 0.6% and (e and f) 0.9%.
in the T60 alloy. The process of forming β phase is slow in a re00 latively low temperature. But as time goes on, most of β phase 0 will transformed into β phase. So the alloy under T60 condition 0 has a high volume fraction of β phase, which has an excellent resistance to the basal slip in the Mg alloy [28]. Therefore, its strengthening effect is developed. By comparison, the alloy under T60 condition can obtain finer and denser precipitates, which will act as the potential obstacles to dislocations. When the samples under loading, the dislocations will pile up in the precipitates, and then prompt to the nucleation of micro-cracks. In this case, the fracture will start from grain 0
interiors, and expand to the grain boundaries. Due to the hindering effect of the grain boundaries, the fracture will be delayed. As a result, the alloy shows a longer fatigue life. Whereas, with low precipitates volume fraction, the dislocations will bypass the precipitates and then pile up at the grain boundaries. Due to the high stress concentration at there, the fracture will be accelerated. With the number of fatigue cycles increasing, the dislocation density gradually increases. And the dislocation entanglement becomes more chaotic, which augments the resistance of subsequent dislocation movement. Moreover, the effect of the solidification microstructure on the cyclic hardening behavior is mainly
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Fig. 8. The optical microstructures of the alloy: (a) T60 and (b) T61.
attributed to the interaction between the dislocation and the precipitates or the grains boundary. When cyclic deformation takes place in the alloy under alternating loading, the cross-slip ability of dislocation will continuously increase [29]. The resistance of dislocation slipping is resulted from the boundaries between the precipitations and the grains, in addition, the refiner grains inevitably shorten the mean free path of dislocation slipping and thus lead to the higher cyclic hardening ability. The Mg-Y-Nd-Zr alloy has refine grains, better grain distributions and the better morphological aspects of phase, so the dislocations easily pile up, entangle and reach to a quasi-stable saturation state in the low strain range.
5. Conclusions In this paper, the low-cycle fatigue properties of the Mg-Y-NdZr alloy by axial cyclic loading under different total strain amplitudes are investigated. And the following conclusions can be drawn from this work: (1) Compared to the heat treatment under T61 condition, the Mg3.5Y-2.4Nd-0.5Zr alloy treated under T60 condition not only could enhance the strength, but also obtain a longer fatigue life.
̅ ⎤⎦ ) and T61 alloy ((c) TEM dark field and (d) SAED pattern with B// ⎡⎣ 1120 ̅ ⎤⎦ ). Fig. 9. TEM images for the T60 alloy ((a) TEM dark field and (b) SAED pattern with B// ⎡⎣ 1120
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(2) During low total strain amplitude (0.3–0.45%), the T60 alloy almost presented cyclic stabilization; and the T61 alloy was increasing hardening until failure. While at higher total strain amplitudes (0.45–0.9%), both them showed continuous hardening. (3) Fatigue cracks of the alloy initiated from the specimen surface, as the strain amplitude increased, the fracture surface became more rough and concavo-convex. (4) The alloy after T60 heat treatment achieved a substantially longer fatigue life than that of the alloy treated under T61 condition at any given total strain amplitude. The improved fatigue resistance of the T60 alloy was mainly attributed to the fine and dense precipitates.
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