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Fatigue cracking behavior of 6063 aluminum alloy for fitting clamps of overhead conductor lines
Materials and Design 88 (2015) 478–484
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Materials and Design journal homepage: www.elsevier.com/locate/jmad
Fatigue cracking behavior of 6063 aluminum alloy for fitting clamps of overhead conductor lines L.R. Zeng a, Z.M. Song a, X.M. Wu b, C.H. Li b, G.P. Zhang a,⁎ a b
Shenyang National Laboratory for Materials Science, Institute of Metal Research, Chinese Academy of Sciences, 72 Wenhua Road, Shenyang 110016, China Institute of Electrical Power Research, Electrical Power Limited Company of Liaoning Province, Shenyang 110006, China
a r t i c l e
i n f o
Article history: Received 4 April 2015 Received in revised form 2 September 2015 Accepted 4 September 2015 Available online 8 September 2015 Keywords: 6063 Al alloy Fatigue Crack growth Stress concentration Stress concentration
a b s t r a c t The 6063 aluminum alloy is a candidate for fitting clamps to fasten and connect the Al alloy overhead conductor lines. The Al alloy fitting clamps with different extents of stress concentration may become quite dangerous under cyclic loading due to the wind-induced swing of the conductor lines. In this study, fatigue properties and cracking behavior of the 6063 aluminum alloy were investigated using three-point-bending specimens with different notch angles. The relationship between the nominal fatigue cracking resistance and the notch angle of the Al alloy specimens was obtained. Fatigue damage and notch angle-dependent cracking behaviors were evaluated. Fatigue reliability of the 6063 alloy specimens with different stress concentration factors was discussed combined with a grain size of the alloy. © 2015 Elsevier Ltd. All rights reserved.
1. Introduction All aluminum alloy conductors (AAACs) have been developed for recent years and gradually replaced the conventional aluminum conductor steel reinforced (ACSR) because the AAACs are lighter than the conventional ACSR. Furthermore, the AAAC has lower breaking loads than the ACSR, especially the use becomes particularly favorable when ice and wind loadings are low. In order to clamp the AAAC, the best way is to select the same material as the AAAC for the fitting clamp. 6063 aluminum alloy, a kind of Al–Mg–Si alloy, is a good candidate not only for the AAAC overhead lines [1,2], but also for fitting clamps to fasten and connect the aluminum alloy conductor lines [3]. Since the Al alloy fitting clamps with different geometrical shapes mechanically connected to the AAAC may have different extents of stress concentration, the stress concentration positions, such as a notch on the clamp, may become quite dangerous under cyclic loading due to the wind-induced swing of the conductor lines. A number of investigations have been conducted to understand the effect of notch on fatigue properties of metals. MacDougall et al. [4] found that the notched specimens exhibited a notch size effect under constant amplitude loading, but the effect was reduced in the long-life region under the variable amplitude load history in 1045 steels. Ren et al. [5] found the similar phenomenon in Udimet 720 alloys. The results showed that the magnitude of notch size effects at the given extent of stress concentration depends on a combination of the plastic zone ⁎ Corresponding author. E-mail address: [email protected] (G.P. Zhang).
http://dx.doi.org/10.1016/j.matdes.2015.09.021 0264-1275/© 2015 Elsevier Ltd. All rights reserved.
size (PZS), maximum plastic strain and stress concentration profile. The stress concentration factor (SCF) is an important factor of the notch effect on the fatigue life. Noda et al. [6–8] studied the effect of different parameters such as the shape, angle and depth of the notch on the SCF. They found that the effect of the notch angle appears to be significant for the shallow notches in the round bar under torsion, and the deep notch bending [6]. Although significant efforts and progress have been made to study the fatigue properties of aluminum alloys for recent years, these studies mainly focus on effects of environment [9–12], microstructure [13–18], fatigue model [12,19,20] and surface treatment [21–24]. In order to evaluate fatigue reliability of the 6063 alloy with different stress concentration extents, in this study we examine fatigue reliability of the 6063 aluminum alloy with different pre-notch angles. Fatigue cracking resistance and fatigue damage behavior of the alloy were evaluated. 2. Experimental procedures The commercial 6063 aluminum alloy for the fitting clamp was selected in this study, which has been subjected to thermal extrusion processing, and the subsequent solid solution treatment. Nominal chemical compositions of the alloy are presented in Table 1. Tensile specimens and three-point-bending (TPB) specimens for fatigue were fabricated by a spark cutting machine. The dimensions of tensile and TPB specimens were schematically illustrated in Fig. 1(a) and (b), respectively. For the TPB specimens, single-edge notches with a depth of 1.7 mm and different angles (θ = 0°, 60°, 90° and 120°) were introduced by the sparking cutting machine. Before the mechanical testing, all the
L.R. Zeng et al. / Materials and Design 88 (2015) 478–484
3. Results and discussion
Table 1 Chemical compositions of 6063 aluminum alloy (wt.%). Mg
Si
Fe
Cu
Ti
Zn
Al
0.45–0.9
0.2–0.6
b0.3
b0.1
b0.1
b0.1
Bal.
specimens were polished mechanically using diamond paste and finally electropolished carefully in a solution of ethylene glycol butyl ether, ethanol and perchloric acid (5:18:2) at a voltage of 18 V and 0 °C. Tensile tests were conducted in a commercial mechanical testing machine (Shimadzu AG-X) at a strain rate of 7.0 × 10−3 s−1. Fatigue tests were performed using MTS® 858 mini Bionix. Stress ratio is 0.1 and loading frequency is 30 Hz. To evaluate and compare fatigue cracking resistance of the TPB specimens with different notch angles, all TPB specimens were loaded cyclically under the same load amplitude until a crack with a length of 0.1 mm initiated from the notch root. The corresponding cycles for the formation of the 0.1 mm-long crack were defined as the nominal fatigue crack initiation life (Ni). The crack lengths were measured by a movable optical microscope. Such a method has been used to evaluate fatigue cracking resistance from a notch in TC11 [25]. The stress intensity factor range (ΔK) at a crack tip for the TPB configuration can be calculated by the applied load range (ΔP) [26]. ΔK ¼
ΔPS BW
3=2
f
479
a ; W
ð1Þ
3.1. Microstructure Fig. 2(a) and (b) presents two typical EBSD inverse pole figures (IPF) of grain orientation distribution in the region ahead of the notch root of the TPB specimens. The alloy consists of equiaxed grains with random orientations. The mean grain size is 93.9 ± 13.3 μm. Radius of the notch root (ρ) was determined based on the SEM observation of the notch root, as presented in Table 2. ρ increases with increasing the notch angle (θ). 3.2. Tensile properties Fig. 3 presents a typical engineering stress–strain curve of the alloy, from which yield and ultimate tensile strengths are determined as 80.85 ± 5.12 MPa and 171.31 ± 6.34 MPa, respectively. Uniform elongation and elongation to fracture are 20.6% and 42.7%, respectively. Obviously, the alloy subjected to the thermal extrusion processing and the subsequent solid solution treatment exhibits a good combination of ductility and strength, as compared with the hard Al lines in the ACSR conductors, which usually have the strength of about 170 MPa and the uniform elongation close to 3% [27]. The above results reveal that the 6063 aluminum alloy clamp has good ductility with uniform elongation up to 20.6%, while the low yield strength of the alloy may facilitate the plastic processing to clamp the conductor lines.
a Þ is a dimensional factor which depends on the ratio of the where f ðW crack length (a) to the width of the specimen (W) and can be expressed by.
f
a W
¼
3ða=W Þ1=2 2½n1 þ 2ða=W Þ½1−ða=W Þ3=2h
1:99−ða=W Þ½1−ða=W Þ 2:15−3:93ða=W Þ þ 2:7ða=W Þ2
io ; ð2Þ
where S is span of the specimen, B is thickness of specimen. After tensile and fatigue tests, all the broken specimens were examined by scanning electron microscopy (SEM, Leo Supra 35).
Fig. 1. Schematic illustration of (a) tensile and (b) three-point-bending fatigue specimens.
Fig. 2. EBSD IPF images of grain structures ahead of the notch root with a notch angle of (a) 0° and (b) 90°.
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Table 2 Radius of notch root with different notch angles. Notch angle, θ (°)
0
60
90
120
Notch root radius, ρ (mm)
0.11
0.17
0.21
0.35
Fig. 4 shows SEM micrographs of surface morphologies on the tensile specimen close to fracture. The extensive slip bands occurred inside grains, leading to incompatible plastic deformation among grains and cracking along either grain boundaries (GBs) or triple junctions, as shown in Fig. 4(b). Fig. 5 presents SEM observations on tensile fracture morphologies of the alloy. Generally, extensive dimples appeared at the fracture surface, indicating the ductile-mode fracture. 3.3. Fatigue cracking resistance and crack growth Fig. 6 shows fatigue crack growth rate (da/dN) of the TPB specimens with different θ as a function of stress intensity factor range (ΔK). In regime A, the fatigue crack growth rates of the specimens with different θ exhibit a large difference, while in regime B (i.e. Paris regime) the fatigue crack growth rates are almost the same for the specimens with different θ. The crack growth rate can be described by the Paris formula given below, da ¼ C ðΔKÞm ; dN
ð3Þ
where C and m are scaling constants. C and m values are obtained as 1.9 × 10−15 m− 3.2·ΜPa−5.47·cycle−1 and 8.4, respectively. Fig. 7 presents a comparison of nominal fatigue cracking resistance (Ni) of specimens with different θ. It is clear that Ni increases with increasing θ. The difference in the growth rate of the physically short crack initiating from the notch root may be influenced by the SCF and the PZS at the notch root. The SCF (kσ) is in turn related to the geometry of the notch depth(c), ρ, θ, the minimum section length (l = W − c) and the c/W ratio [28] (Fig. 1). kσ can be estimated by the below formula [8,29]. ðkσs −1Þðkσd −1Þ ffi; kσ ¼ 1 þ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 ðkσs −1Þ þ ðkσ d −1Þ
ð4Þ
where the parameters kσd, kσe and kσs are given as, 2 kσd ¼
. l
ρ
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi . l þ 1 −β1 þ1
4 β2
. l
ρ
ρ
þ 1 −3β1
ð5Þ
Fig. 4. SEM micrographs of deformation morphologies at the tensile specimen surfaces observed at (a) low magnification and (b) high magnification on the dash-line region in (a). The loading direction is along the horizontal direction.
kσs ¼ 1:035 þ 0:0261η−0:1451η2 þ 0:0842η3 kσe
ð6Þ
qffiffiffiffiffiffiffi kσe ¼ 1:121−0:2846η þ 0:3397η2 −0:1544η3 1 þ 2 c =ρ
ð7Þ
rffiffiffiffiffiffiffi .ffi l þ1 ρ ρ rffiffiffiffiffiffiffi β1 ¼ . . ffi rffiffiffiffiffiffiffi .ffi l þ 1 tan−1 l þ l 2
. l
ρ
ρ
ð8Þ
ρ
3 = 2 4 l ρ r. β 2 ¼ rffiffiffiffiffiffiffi .ffi . l 3 þ l −1 tan−1 l ρ
qffiffiffiffiffiffiffi η ¼ ρ =c :
Fig. 3. Tensile engineering stress-engineering strain curves of 6063 aluminum alloy.
ρ
ð9Þ
ρ
ð10Þ
Using Eqs. (4)–(10) and parameters shown in Fig. 1 and Table 2, the SCF was calculated and shown in Fig. 8, which demonstrates an evident decrease with increasing θ. As the concentration stress ahead of the notch root is higher than the tensile yield strength, the region at the notch root is deformed plastically. Then, a small plastic zone with a size of rpl appears at the notch root. In the plastic zone, a compressive residual stress leads to the plastically-
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481
Fig. 7. Relation between nominal fatigue crack initiation life and notch angle.
of generally increased crack propagation in the corrosive medium. The notch PZS can be estimated by [28] π−θ r pl ¼ ρ exp −1 : 2
Fig. 5. SEM images of tensile fracture morphologies of 6063 aluminum alloy observed at (a) low and (b) high magnifications.
induced crack closure, thus effectively decreases ΔK at the crack tip and crack propagation rate [30]. Zamponi et al. [9] found that the effect of the larger plastic zone partly compensates the otherwise negative effect
Fig. 6. Fatigue crack growth rate (da/dN) of 6063 alloys as a function of stress intensity factor range (ΔK).
ð11Þ
Using Eq. (11), the PZS was estimated and also presented in Fig. 8. It is clear that the larger the notch angle, the smaller the plastic zone size. When a fatigue crack initiates from the notch root and grows, there are two factors controlling the development of the physically short crack, i.e. SCF and crack closure. At the same applied load, the larger SCF generates higher stresses acting on the {111} slip planes within grains ahead of the notch root, and leads to the larger notch plastic zone where more irreversible slip damage forms. Thus, crack initiation would take place preferentially from the sharp notch root with higher SCF. On the contrary, the crack closure, which can be induced by the crack surface roughness and plastic zone ahead of the crack, provides a resistance to the fatigue crack growth. McClung et al. [31] suggested that a variety of closure mechanisms may be operative in the near threshold region, including residual plasticity, fracture surface roughness, and fracture surface oxide. Parry et al. [32] had the similar conclusion using finite element modeling. In the present case for the same material and the applied load, it is expected that the four specimens with different θ are subjected to the roughness-induced crack closure when a short crack with the same length initiates from the notches. Furthermore, the plastic-induced crack closure at the notch root also plays an important role in retarding fatigue cracking. Based on Eq. (11), the
Fig. 8. Variations of stress concentration factor (SCF) and plastic zone size (PZS) as a function of the notch angle.
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PZS at the notch root with different θ is quite different and larger than the average grain size. Thus, the short crack (0.1 mm) is actually surrounded by the notch plastic zone. It is believed that both the notch PZS and the plastic-induced crack closure generated in the plastic zone influences fatigue crack growth. Considering both contrary effects caused by SCF and the crack closure, the evidently low fatigue cracking resistance of the θ = 0° specimen shown in Fig. 7 reveals that the detrimental effect of SCF generated at the sharp notch root overwhelms the beneficial effect of the plastic zone as θ decreasing and dominates the fatigue cracking ability. Once the physically short fatigue crack grows over the notch plastic zone, the effect of the notch angle on the SCF is getting weak and the PZS ahead of the crack tip may become dominated. In this case, the crack growth rate for the specimen with different θ is the same in the regime B shown in Fig. 6. 3.4. Fatigue cracking behavior Fatigue cracking behaviors in all the specimens with different θ were observed, and they show a similar behavior. Thus, Fig. 9(a) only shows a typical SEM micrograph of cracking behavior of the θ = 0° specimen. During initial cracking from the notch root (zone 1 in Fig. 9(a)), the crack mainly grew along crystallographic slip planes in a zig–zag mode, as shown in Fig. 9(b). Fig. 9(c) shows some parallel slip lines appearing in two grains. The EBSD IPF image shows that the crack
grew along (1-11) and (1-1-1) planes alternatively, indicating that the fatigue damage has formed along the crystallographic slip planes {111} plane (see Fig. 9(d)). With the crack further advancing (zone 2 in Fig. 9(a)), the surface cracking along GBs became evident, as indicated by an arrow in Fig. 9(e). The crack deviated from the crystallographic slip planes, and turned into GB cracking, especially in the region ahead of the crack tip. Meanwhile, the secondary cracks became more evident, as shown in Fig. 9(e). With further increasing the crack length (zone 3 in Fig. 9(a)), large plastic deformation has formed, as characterized by extensive slip lines at the crack tip, as shown in Fig. 9(f). Fig. 10(a) shows an SEM image of fatigue fractograph of the θ = 0° specimen. Fig. 10(b)–(f) shows detailed fracture features in regions I, II and III, respectively. In region I (Fig. 10(b)) close to the notch root (lower ΔK), there are many smooth facets crossing the whole grains, which are corresponding to the crystallographic slip planes. This led to a locally rough fracture surface. Fig. 10(c) shows a close observation on the facets. In region II in the center of the specimen (Fig. 10(d)) corresponding to the stable stage of crack growth (higher ΔK in regime B in Fig. 6), there are a lot of fatigue striations and secondary cracks, as indicated by arrows in Fig. 10(d) and (e). The crack advanced in transgranular mode. In region III (Fig. 10(f)) close to the specimen surface corresponding to stable crack growth stage (regime B in Fig. 6), clear GB cracking happened, which is consistent with the observations in Fig. 10(e).
Fig. 9. SEM observations on fatigue cracking behaviors in the θ = 0° specimen of 6063 aluminum alloy, (a) overview, (b) and (c) close observations on region 1 in (a), and (d) EBSD image on the region in (c). (e) and (f) close observations for zones 2 and 3 in (a).
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483
Fig. 10. Fatigue fracture morphologies of the 6063 alloy, (a) overview, (b) and (c) close observations for zone I in (a), (d) and (e) close observation for zone II, (f) close observation for zone III.
In general, the notched specimens show the ΔK-dependent cracking behavior, i.e. the crack mainly propagated along the crystallographic slip planes at lower ΔK following a typical stage I crack growth, and then that gradually changed from the crystallographic mode to the transgranular mode with further increasing ΔK. This may lead to a difference in the crack growth rate with varying θ at the regime A, but the difference disappears at the regime B, as shown in Fig. 6. Microscopically, a comparison between the PZS (Fig. 8) and the mean grain size (Fig. 2) indicates that there are only several grains within the plastic zone at the notch root, thus cyclic plastic deformation may happen in the grains and the slip band-induced cracking is the primary mechanism for crack initiation (Fig. 9(c)). This is evidently shown by the SEM observations and the EBSD characterization (Fig. 9(b)–(d)). The present results indicate that a decrease in the grain size of the 6063 alloy for the fitting clamp may not only enhance the yield strength of the alloy and thus promote the resistance of crack initiation from the place with high stress concentration, but also may increase the number of grains in plastic zone at the notch root, which is expected to provide more resistance to fatigue cracking because more GBs would take part in hindering the crack growth. 4. Conclusions Fatigue properties and cracking behavior of the 6063 aluminum alloy for fitting clamps of overhead conductor lines were examined using TPB specimens with different notch angles. The following conclusions can be drawn. (1) The nominal fatigue cracking resistance from the notch of the 6063 alloy increases with increasing the notch angle from 0° to 120°. Stress concentration factor may dominate the extent of fatigue cracking damage of the 6063 alloy. (2) With increasing the crack length, the fatigue cracking mode of the 6063 alloy may change from the slip band-induced crystallographic mode along slip planes at the notch root to the transgranular mode. (3) Fatigue crack growth rate of the 6063 alloy in the region close to the notch root with different notch angles is quite different, while that became independent on the notch angle as the crack becomes longer.
Acknowledgments This work was supported by the National Natural Science Foundation of China (NSFC, Grant No. 51371180).
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Contents lists available at ScienceDirect
Materials and Design journal homepage: www.elsevier.com/locate/jmad
Fatigue cracking behavior of 6063 aluminum alloy for fitting clamps of overhead conductor lines L.R. Zeng a, Z.M. Song a, X.M. Wu b, C.H. Li b, G.P. Zhang a,⁎ a b
Shenyang National Laboratory for Materials Science, Institute of Metal Research, Chinese Academy of Sciences, 72 Wenhua Road, Shenyang 110016, China Institute of Electrical Power Research, Electrical Power Limited Company of Liaoning Province, Shenyang 110006, China
a r t i c l e
i n f o
Article history: Received 4 April 2015 Received in revised form 2 September 2015 Accepted 4 September 2015 Available online 8 September 2015 Keywords: 6063 Al alloy Fatigue Crack growth Stress concentration Stress concentration
a b s t r a c t The 6063 aluminum alloy is a candidate for fitting clamps to fasten and connect the Al alloy overhead conductor lines. The Al alloy fitting clamps with different extents of stress concentration may become quite dangerous under cyclic loading due to the wind-induced swing of the conductor lines. In this study, fatigue properties and cracking behavior of the 6063 aluminum alloy were investigated using three-point-bending specimens with different notch angles. The relationship between the nominal fatigue cracking resistance and the notch angle of the Al alloy specimens was obtained. Fatigue damage and notch angle-dependent cracking behaviors were evaluated. Fatigue reliability of the 6063 alloy specimens with different stress concentration factors was discussed combined with a grain size of the alloy. © 2015 Elsevier Ltd. All rights reserved.
1. Introduction All aluminum alloy conductors (AAACs) have been developed for recent years and gradually replaced the conventional aluminum conductor steel reinforced (ACSR) because the AAACs are lighter than the conventional ACSR. Furthermore, the AAAC has lower breaking loads than the ACSR, especially the use becomes particularly favorable when ice and wind loadings are low. In order to clamp the AAAC, the best way is to select the same material as the AAAC for the fitting clamp. 6063 aluminum alloy, a kind of Al–Mg–Si alloy, is a good candidate not only for the AAAC overhead lines [1,2], but also for fitting clamps to fasten and connect the aluminum alloy conductor lines [3]. Since the Al alloy fitting clamps with different geometrical shapes mechanically connected to the AAAC may have different extents of stress concentration, the stress concentration positions, such as a notch on the clamp, may become quite dangerous under cyclic loading due to the wind-induced swing of the conductor lines. A number of investigations have been conducted to understand the effect of notch on fatigue properties of metals. MacDougall et al. [4] found that the notched specimens exhibited a notch size effect under constant amplitude loading, but the effect was reduced in the long-life region under the variable amplitude load history in 1045 steels. Ren et al. [5] found the similar phenomenon in Udimet 720 alloys. The results showed that the magnitude of notch size effects at the given extent of stress concentration depends on a combination of the plastic zone ⁎ Corresponding author. E-mail address: [email protected] (G.P. Zhang).
http://dx.doi.org/10.1016/j.matdes.2015.09.021 0264-1275/© 2015 Elsevier Ltd. All rights reserved.
size (PZS), maximum plastic strain and stress concentration profile. The stress concentration factor (SCF) is an important factor of the notch effect on the fatigue life. Noda et al. [6–8] studied the effect of different parameters such as the shape, angle and depth of the notch on the SCF. They found that the effect of the notch angle appears to be significant for the shallow notches in the round bar under torsion, and the deep notch bending [6]. Although significant efforts and progress have been made to study the fatigue properties of aluminum alloys for recent years, these studies mainly focus on effects of environment [9–12], microstructure [13–18], fatigue model [12,19,20] and surface treatment [21–24]. In order to evaluate fatigue reliability of the 6063 alloy with different stress concentration extents, in this study we examine fatigue reliability of the 6063 aluminum alloy with different pre-notch angles. Fatigue cracking resistance and fatigue damage behavior of the alloy were evaluated. 2. Experimental procedures The commercial 6063 aluminum alloy for the fitting clamp was selected in this study, which has been subjected to thermal extrusion processing, and the subsequent solid solution treatment. Nominal chemical compositions of the alloy are presented in Table 1. Tensile specimens and three-point-bending (TPB) specimens for fatigue were fabricated by a spark cutting machine. The dimensions of tensile and TPB specimens were schematically illustrated in Fig. 1(a) and (b), respectively. For the TPB specimens, single-edge notches with a depth of 1.7 mm and different angles (θ = 0°, 60°, 90° and 120°) were introduced by the sparking cutting machine. Before the mechanical testing, all the
L.R. Zeng et al. / Materials and Design 88 (2015) 478–484
3. Results and discussion
Table 1 Chemical compositions of 6063 aluminum alloy (wt.%). Mg
Si
Fe
Cu
Ti
Zn
Al
0.45–0.9
0.2–0.6
b0.3
b0.1
b0.1
b0.1
Bal.
specimens were polished mechanically using diamond paste and finally electropolished carefully in a solution of ethylene glycol butyl ether, ethanol and perchloric acid (5:18:2) at a voltage of 18 V and 0 °C. Tensile tests were conducted in a commercial mechanical testing machine (Shimadzu AG-X) at a strain rate of 7.0 × 10−3 s−1. Fatigue tests were performed using MTS® 858 mini Bionix. Stress ratio is 0.1 and loading frequency is 30 Hz. To evaluate and compare fatigue cracking resistance of the TPB specimens with different notch angles, all TPB specimens were loaded cyclically under the same load amplitude until a crack with a length of 0.1 mm initiated from the notch root. The corresponding cycles for the formation of the 0.1 mm-long crack were defined as the nominal fatigue crack initiation life (Ni). The crack lengths were measured by a movable optical microscope. Such a method has been used to evaluate fatigue cracking resistance from a notch in TC11 [25]. The stress intensity factor range (ΔK) at a crack tip for the TPB configuration can be calculated by the applied load range (ΔP) [26]. ΔK ¼
ΔPS BW
3=2
f
479
a ; W
ð1Þ
3.1. Microstructure Fig. 2(a) and (b) presents two typical EBSD inverse pole figures (IPF) of grain orientation distribution in the region ahead of the notch root of the TPB specimens. The alloy consists of equiaxed grains with random orientations. The mean grain size is 93.9 ± 13.3 μm. Radius of the notch root (ρ) was determined based on the SEM observation of the notch root, as presented in Table 2. ρ increases with increasing the notch angle (θ). 3.2. Tensile properties Fig. 3 presents a typical engineering stress–strain curve of the alloy, from which yield and ultimate tensile strengths are determined as 80.85 ± 5.12 MPa and 171.31 ± 6.34 MPa, respectively. Uniform elongation and elongation to fracture are 20.6% and 42.7%, respectively. Obviously, the alloy subjected to the thermal extrusion processing and the subsequent solid solution treatment exhibits a good combination of ductility and strength, as compared with the hard Al lines in the ACSR conductors, which usually have the strength of about 170 MPa and the uniform elongation close to 3% [27]. The above results reveal that the 6063 aluminum alloy clamp has good ductility with uniform elongation up to 20.6%, while the low yield strength of the alloy may facilitate the plastic processing to clamp the conductor lines.
a Þ is a dimensional factor which depends on the ratio of the where f ðW crack length (a) to the width of the specimen (W) and can be expressed by.
f
a W
¼
3ða=W Þ1=2 2½n1 þ 2ða=W Þ½1−ða=W Þ3=2h
1:99−ða=W Þ½1−ða=W Þ 2:15−3:93ða=W Þ þ 2:7ða=W Þ2
io ; ð2Þ
where S is span of the specimen, B is thickness of specimen. After tensile and fatigue tests, all the broken specimens were examined by scanning electron microscopy (SEM, Leo Supra 35).
Fig. 1. Schematic illustration of (a) tensile and (b) three-point-bending fatigue specimens.
Fig. 2. EBSD IPF images of grain structures ahead of the notch root with a notch angle of (a) 0° and (b) 90°.
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Table 2 Radius of notch root with different notch angles. Notch angle, θ (°)
0
60
90
120
Notch root radius, ρ (mm)
0.11
0.17
0.21
0.35
Fig. 4 shows SEM micrographs of surface morphologies on the tensile specimen close to fracture. The extensive slip bands occurred inside grains, leading to incompatible plastic deformation among grains and cracking along either grain boundaries (GBs) or triple junctions, as shown in Fig. 4(b). Fig. 5 presents SEM observations on tensile fracture morphologies of the alloy. Generally, extensive dimples appeared at the fracture surface, indicating the ductile-mode fracture. 3.3. Fatigue cracking resistance and crack growth Fig. 6 shows fatigue crack growth rate (da/dN) of the TPB specimens with different θ as a function of stress intensity factor range (ΔK). In regime A, the fatigue crack growth rates of the specimens with different θ exhibit a large difference, while in regime B (i.e. Paris regime) the fatigue crack growth rates are almost the same for the specimens with different θ. The crack growth rate can be described by the Paris formula given below, da ¼ C ðΔKÞm ; dN
ð3Þ
where C and m are scaling constants. C and m values are obtained as 1.9 × 10−15 m− 3.2·ΜPa−5.47·cycle−1 and 8.4, respectively. Fig. 7 presents a comparison of nominal fatigue cracking resistance (Ni) of specimens with different θ. It is clear that Ni increases with increasing θ. The difference in the growth rate of the physically short crack initiating from the notch root may be influenced by the SCF and the PZS at the notch root. The SCF (kσ) is in turn related to the geometry of the notch depth(c), ρ, θ, the minimum section length (l = W − c) and the c/W ratio [28] (Fig. 1). kσ can be estimated by the below formula [8,29]. ðkσs −1Þðkσd −1Þ ffi; kσ ¼ 1 þ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 ðkσs −1Þ þ ðkσ d −1Þ
ð4Þ
where the parameters kσd, kσe and kσs are given as, 2 kσd ¼
. l
ρ
sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi . l þ 1 −β1 þ1
4 β2
. l
ρ
ρ
þ 1 −3β1
ð5Þ
Fig. 4. SEM micrographs of deformation morphologies at the tensile specimen surfaces observed at (a) low magnification and (b) high magnification on the dash-line region in (a). The loading direction is along the horizontal direction.
kσs ¼ 1:035 þ 0:0261η−0:1451η2 þ 0:0842η3 kσe
ð6Þ
qffiffiffiffiffiffiffi kσe ¼ 1:121−0:2846η þ 0:3397η2 −0:1544η3 1 þ 2 c =ρ
ð7Þ
rffiffiffiffiffiffiffi .ffi l þ1 ρ ρ rffiffiffiffiffiffiffi β1 ¼ . . ffi rffiffiffiffiffiffiffi .ffi l þ 1 tan−1 l þ l 2
. l
ρ
ρ
ð8Þ
ρ
3 = 2 4 l ρ r. β 2 ¼ rffiffiffiffiffiffiffi .ffi . l 3 þ l −1 tan−1 l ρ
qffiffiffiffiffiffiffi η ¼ ρ =c :
Fig. 3. Tensile engineering stress-engineering strain curves of 6063 aluminum alloy.
ρ
ð9Þ
ρ
ð10Þ
Using Eqs. (4)–(10) and parameters shown in Fig. 1 and Table 2, the SCF was calculated and shown in Fig. 8, which demonstrates an evident decrease with increasing θ. As the concentration stress ahead of the notch root is higher than the tensile yield strength, the region at the notch root is deformed plastically. Then, a small plastic zone with a size of rpl appears at the notch root. In the plastic zone, a compressive residual stress leads to the plastically-
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Fig. 7. Relation between nominal fatigue crack initiation life and notch angle.
of generally increased crack propagation in the corrosive medium. The notch PZS can be estimated by [28] π−θ r pl ¼ ρ exp −1 : 2
Fig. 5. SEM images of tensile fracture morphologies of 6063 aluminum alloy observed at (a) low and (b) high magnifications.
induced crack closure, thus effectively decreases ΔK at the crack tip and crack propagation rate [30]. Zamponi et al. [9] found that the effect of the larger plastic zone partly compensates the otherwise negative effect
Fig. 6. Fatigue crack growth rate (da/dN) of 6063 alloys as a function of stress intensity factor range (ΔK).
ð11Þ
Using Eq. (11), the PZS was estimated and also presented in Fig. 8. It is clear that the larger the notch angle, the smaller the plastic zone size. When a fatigue crack initiates from the notch root and grows, there are two factors controlling the development of the physically short crack, i.e. SCF and crack closure. At the same applied load, the larger SCF generates higher stresses acting on the {111} slip planes within grains ahead of the notch root, and leads to the larger notch plastic zone where more irreversible slip damage forms. Thus, crack initiation would take place preferentially from the sharp notch root with higher SCF. On the contrary, the crack closure, which can be induced by the crack surface roughness and plastic zone ahead of the crack, provides a resistance to the fatigue crack growth. McClung et al. [31] suggested that a variety of closure mechanisms may be operative in the near threshold region, including residual plasticity, fracture surface roughness, and fracture surface oxide. Parry et al. [32] had the similar conclusion using finite element modeling. In the present case for the same material and the applied load, it is expected that the four specimens with different θ are subjected to the roughness-induced crack closure when a short crack with the same length initiates from the notches. Furthermore, the plastic-induced crack closure at the notch root also plays an important role in retarding fatigue cracking. Based on Eq. (11), the
Fig. 8. Variations of stress concentration factor (SCF) and plastic zone size (PZS) as a function of the notch angle.
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PZS at the notch root with different θ is quite different and larger than the average grain size. Thus, the short crack (0.1 mm) is actually surrounded by the notch plastic zone. It is believed that both the notch PZS and the plastic-induced crack closure generated in the plastic zone influences fatigue crack growth. Considering both contrary effects caused by SCF and the crack closure, the evidently low fatigue cracking resistance of the θ = 0° specimen shown in Fig. 7 reveals that the detrimental effect of SCF generated at the sharp notch root overwhelms the beneficial effect of the plastic zone as θ decreasing and dominates the fatigue cracking ability. Once the physically short fatigue crack grows over the notch plastic zone, the effect of the notch angle on the SCF is getting weak and the PZS ahead of the crack tip may become dominated. In this case, the crack growth rate for the specimen with different θ is the same in the regime B shown in Fig. 6. 3.4. Fatigue cracking behavior Fatigue cracking behaviors in all the specimens with different θ were observed, and they show a similar behavior. Thus, Fig. 9(a) only shows a typical SEM micrograph of cracking behavior of the θ = 0° specimen. During initial cracking from the notch root (zone 1 in Fig. 9(a)), the crack mainly grew along crystallographic slip planes in a zig–zag mode, as shown in Fig. 9(b). Fig. 9(c) shows some parallel slip lines appearing in two grains. The EBSD IPF image shows that the crack
grew along (1-11) and (1-1-1) planes alternatively, indicating that the fatigue damage has formed along the crystallographic slip planes {111} plane (see Fig. 9(d)). With the crack further advancing (zone 2 in Fig. 9(a)), the surface cracking along GBs became evident, as indicated by an arrow in Fig. 9(e). The crack deviated from the crystallographic slip planes, and turned into GB cracking, especially in the region ahead of the crack tip. Meanwhile, the secondary cracks became more evident, as shown in Fig. 9(e). With further increasing the crack length (zone 3 in Fig. 9(a)), large plastic deformation has formed, as characterized by extensive slip lines at the crack tip, as shown in Fig. 9(f). Fig. 10(a) shows an SEM image of fatigue fractograph of the θ = 0° specimen. Fig. 10(b)–(f) shows detailed fracture features in regions I, II and III, respectively. In region I (Fig. 10(b)) close to the notch root (lower ΔK), there are many smooth facets crossing the whole grains, which are corresponding to the crystallographic slip planes. This led to a locally rough fracture surface. Fig. 10(c) shows a close observation on the facets. In region II in the center of the specimen (Fig. 10(d)) corresponding to the stable stage of crack growth (higher ΔK in regime B in Fig. 6), there are a lot of fatigue striations and secondary cracks, as indicated by arrows in Fig. 10(d) and (e). The crack advanced in transgranular mode. In region III (Fig. 10(f)) close to the specimen surface corresponding to stable crack growth stage (regime B in Fig. 6), clear GB cracking happened, which is consistent with the observations in Fig. 10(e).
Fig. 9. SEM observations on fatigue cracking behaviors in the θ = 0° specimen of 6063 aluminum alloy, (a) overview, (b) and (c) close observations on region 1 in (a), and (d) EBSD image on the region in (c). (e) and (f) close observations for zones 2 and 3 in (a).
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Fig. 10. Fatigue fracture morphologies of the 6063 alloy, (a) overview, (b) and (c) close observations for zone I in (a), (d) and (e) close observation for zone II, (f) close observation for zone III.
In general, the notched specimens show the ΔK-dependent cracking behavior, i.e. the crack mainly propagated along the crystallographic slip planes at lower ΔK following a typical stage I crack growth, and then that gradually changed from the crystallographic mode to the transgranular mode with further increasing ΔK. This may lead to a difference in the crack growth rate with varying θ at the regime A, but the difference disappears at the regime B, as shown in Fig. 6. Microscopically, a comparison between the PZS (Fig. 8) and the mean grain size (Fig. 2) indicates that there are only several grains within the plastic zone at the notch root, thus cyclic plastic deformation may happen in the grains and the slip band-induced cracking is the primary mechanism for crack initiation (Fig. 9(c)). This is evidently shown by the SEM observations and the EBSD characterization (Fig. 9(b)–(d)). The present results indicate that a decrease in the grain size of the 6063 alloy for the fitting clamp may not only enhance the yield strength of the alloy and thus promote the resistance of crack initiation from the place with high stress concentration, but also may increase the number of grains in plastic zone at the notch root, which is expected to provide more resistance to fatigue cracking because more GBs would take part in hindering the crack growth. 4. Conclusions Fatigue properties and cracking behavior of the 6063 aluminum alloy for fitting clamps of overhead conductor lines were examined using TPB specimens with different notch angles. The following conclusions can be drawn. (1) The nominal fatigue cracking resistance from the notch of the 6063 alloy increases with increasing the notch angle from 0° to 120°. Stress concentration factor may dominate the extent of fatigue cracking damage of the 6063 alloy. (2) With increasing the crack length, the fatigue cracking mode of the 6063 alloy may change from the slip band-induced crystallographic mode along slip planes at the notch root to the transgranular mode. (3) Fatigue crack growth rate of the 6063 alloy in the region close to the notch root with different notch angles is quite different, while that became independent on the notch angle as the crack becomes longer.
Acknowledgments This work was supported by the National Natural Science Foundation of China (NSFC, Grant No. 51371180).
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