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Chapter 15 Effects of mercury on glutathione and glutathione-dependent enzymes in catfish (Ictalurus melas R.)
Materials Science & Engineering A 625 (2015) 369–373
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Materials Science & Engineering A journal homepage: www.elsevier.com/locate/msea
Improving the tensile strength of Mg–7Sn–1Al–1Zn alloy through artificial cooling during extrusion Sung Hyuk Park n, Ha Sik Kim, Bong Sun You Light Metal Division, Korea Institute of Materials Science (KIMS), Changwon 642-831, Republic of Korea
art ic l e i nf o
a b s t r a c t
Article history: Received 6 November 2014 Received in revised form 27 November 2014 Accepted 2 December 2014 Available online 10 December 2014
An Mg–7Sn–1Al–1Zn alloy known to have excellent extrudability and superior strength was subjected to artificial cooling during indirect extrusion by directly spraying water onto the extruded rod at the die exit. The results obtained revealed that this artificial cooling dramatically reduces the temperature of the deformation zone during extrusion, thereby creating a finer grain size, an intensified texture and a greater amount of precipitates when compared to extrusion without artificial cooling. The yield and tensile strength of the extruded alloy is also significantly improved, which is attributed to the effects of grain refinement in combination with an enhanced texture and precipitate hardening. & 2014 Elsevier B.V. All rights reserved.
Keywords: Magnesium alloy Extrusion Artificial cooling Tensile strength Precipitation
1. Introduction The automotive industry's demand for wrought magnesium (Mg) alloys has been increasing in recent years due to their mechanical superiority over comparable cast alloys, not to mention their extremely low density relative to other metallic materials. In the case of extruded Mg alloys, however, supply and application has so far been greatly limited by the low maximum extrusion speed of commercial high-strength Mg alloys such as AZ80 and ZK60 [1]; a slower manufacturing speed invariably means a higher final product cost. This has led to an increased interest in Mg–Sn based alloys due to their ability to be extruded without developing surface cracks at speeds equivalent to those used for aluminum alloys [2], which is attributed to the formation of a thermally stable Mg2Sn phase in place of low-melting temperature Mg17Al12 or MgZn2 phases [2–4]. Of all the Mg–Sn based alloys, those with a high Sn content have been the subject of particular focus by virtue of the very attractive combination of high strength and superior extrudability that they offer. Indeed, Elsayed et al. [5] have already developed an Mg–9.8Sn–3.0Al– 0.5Zn (wt%) (TAZ1031) alloy exhibiting a symmetrical tensioncompression yield strength and high ultimate tensile strength (UTS) of 358 MPa in an as-extruded state, while Sasaki et al. [4] have created an Mg–9.8Sn–1.2Zn–1.0Al (wt%) alloy with a high tensile yield strength (TYS) of 308 MPa through extrusion at a low temperature of 250 1C with a low ram speed of 0.1 mm/s.
n
Corresponding author. Tel.: þ 82 55 280 3516; fax: þ 82 55 280 3599. E-mail address: [email protected] (S.H. Park).
http://dx.doi.org/10.1016/j.msea.2014.12.011 0921-5093/& 2014 Elsevier B.V. All rights reserved.
Meanwhile, Park et al. [6] have also demonstrated an extruded Mg–7.95Sn–0.95Al–0.95Zn (wt%) (TAZ811) alloy exhibiting superior tensile and compressive strengths compared to extruded AZ31 alloy. In our previous study [7], an extraordinarily high-strength TAZ811 alloy (TYS of 390 MPa, UTS of 405 MPa) was fabricated by extruding a billet subjected to cold forging prior to extrusion. It has also been recently reported that Mg–7Sn–1Al–1Zn (wt%) (TAZ711) alloy has comparable tensile strength to commercial high-strength AZ80 alloy extruded under the same conditions [2], in addition to an excellent extrudability that is demonstrable by a maximum extrusion speed of more than 12 m/min [2,8]. It is well known that the microstructure and mechanical properties of extruded Mg alloys are strongly dependent on extrusion parameters such as billet size, die geometry, reduction ratio, initial temperature, and ram speed [8–14]. For instance, with an increase in extrusion speed, the tensile strength of the extruded alloy is gradually reduced by an increase in the final grain size. Similarly, although Mg–Sn based alloys can be rapidly extruded without cracking, such high-speed extrusion invariably induces some degree of deformation heating that reduces the strength of the extruded alloy through grain growth [8,14]. Cheng et al. [14] have also shown a decrease in the tensile strength of extruded TAZ811 alloy (from 244 to 199 MPa in TYS and from 312 to 286 MPa in UTS) with increasing extrusion exit speeds (from 2 to 10 m/min). The additional heating caused by abrupt plastic deformation and friction during high-speed extrusion should therefore be reduced in order to limit grain growth and ensure high strength in the extruded alloy. In this regard, one method that has proven effective in suppressing excessive temperature rise at high extrusion speeds is the application of artificial cooling,
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S.H. Park et al. / Materials Science & Engineering A 625 (2015) 369–373
with Kim et al. [15] using this to achieve an increase in the tensile strength of indirectly extruded AZ31 alloy. However, there has as yet been no investigation into the effect of artificial cooling on Mg–Sn based alloys developed specifically for high speed extrusion. Moreover, although it is well known that the fine precipitates generated during extrusion play an important role in determining the strength of the final product, the influence of artificial cooling on dynamic precipitation has been overlooked by previous studies concerned only with AZ31 alloy that contains no second phases. The main objective of the present study is therefore to determine the microstructural variation that induced by artificial cooling in terms of grain size, texture and precipitate in TAZ711 that is known to have a superior strength and excellent extrudability. The effect of this variation on the tensile properties of the extruded alloy is also explored and discussed herein.
The tensile properties of the extruded samples were measured at room temperature using an Instron 4206 universal testing machine with a strain rate of 1.0 10 3 s 1. Dog-bone shaped (gage section: ∅6 mm 25 mm) specimens were used for tensile testing, the axes of which corresponded to the extrusion direction (ED). The tensile test for each sample was repeated three times to ensure repeatability and confirm the consistency of the results, but for the sake of simplicity, a representative curve for each test was used. It should also be noted that the 6 mm gage diameter of the tensile specimens represents the center of the extruded bars (∅11.3 mm), and thus their tensile properties are more closely related to the microstructure within this region than that of the sub-surface region. This means that the microstructural characteristics of the extruded bars were analyzed with a particular focus on their center region.
2. Experimental procedure
3. Results and discussion
Cast billets of TAZ711 alloy were prepared by first melting in an electric resistance furnace at 800 1C under an inert atmosphere containing a mixture of CO2 and SF6, and then pouring into a steel mold pre-heated to 210 1C. The chemical composition of the resulting billet was found through inductively couple plasma spectrometry to be 6.81Sn–1.10Al–1.07Zn (wt%). These billets were homogenized at 500 1C for 24 h to fully dissolve any macrosegregation of the solute elements and intermetallic compounds formed during solidification, and then water-quenched to obtain a supersaturated solid-solution containing no Mg2Sn phase. Using a previously described process [16], two billets (∅50 mm 150 mm) pre-heated to 350 1C were indirectly extruded at a temperature of 350 1C using a ram speed of 1 mm/s and an extrusion ratio of 20, with artificial cooling being applied to only one of these. A schematic diagram of the indirect extrusion process and associated artificial cooling system is shown in Fig. 1a. In this, cold water was fed at a rate of 3 ℓ/min through two inlets located on the surface of the stem, being then transferred through holes inside the stem and die to spray directly onto the extruded rod at the die exit. The variation in die temperature during extrusion was measured using a thermocouple installed inside the die (Fig. 1a). The microstructural and textural characteristics of the extruded samples were analyzed by optical microscopy (OM), electron backscatter diffraction (EBSD), field emission scanning electron microscopy (FE-SEM), and X-ray diffraction (XRD). The area fraction of recrystallized grains and secondary phases in each of the extruded samples was obtained by averaging the values taken from five low-magnification OM and SEM images, respectively.
Fig. 1b shows the measured variation in the extrusion load and die temperature during extrusion, the latter being considered a more accurate reflection of the actual temperature applied to the alloy during extrusion. In the case of extrusion without cooling, although the initial extrusion temperature is 350 1C, the die temperature gradually rises to 378 1C during the initial stage of extrusion due to the heat generated by plastic deformation and friction. This reduces the hot working flow stress, and subsequently leads to a decrease in the extrusion load. However, as the extrusion process stabilizes, the die temperature decreases slightly to 370 1C; a temperature that is maintained to nearly the end of extrusion. Conversely, when artificial cooling is applied during extrusion, the die temperature is rapidly reduced to 260 1C immediately after the extruded bar exits the die, this being despite the corresponding increase in extrusion load due to the greater flow stress of the billet. Furthermore, this reduced temperature remains essentially unchanged throughout whole extrusion process, which indicates that artificial cooling applied directly to the extruded bar significantly reduces its actual temperature during extrusion from 370 to 260 1C (by 110 1C). This, in turn, should result in a significant grain refinement through the suppression of grain growth. As confirmed in our previous study [17], homogenized TAZ711 alloy consists of equiaxed grains with an average size of 340 um and a very small amount of Al3Fe or Al8Mn5 inclusions; however, any coarse Mg2Sn phases formed during solidification are fully dissolved by homogenization treatment. Fig. 2 shows the microstructure and texture within the center region of the artificially cooled (AC) and naturally cooled (NC) samples of extruded TAZ711
140
600
without cooling with cooling
550
120
Die
Billet
Stem Extrudate
Die
Stem Inlet
500
Extrusion load
450
100
400
80
350
Die temperature
300
60
250
Ex xtrusion load (ton)
90o
Thermocouple Inlet
Die temperature (oC C)
Sealing element
Container
40
200 0
20
40
60 100 80 Ram distance (mm)
120
Fig. 1. (a) Schematic diagram of the indirect extrusion process and artificial cooling system used in this study. (b) Variation in die temperature and extrusion load during extrusion.
S.H. Park et al. / Materials Science & Engineering A 625 (2015) 369–373
fDRX = 73%
dDRX = 5.4 µm
ED
ED
371
Max.=4.0 non-DRXed DRXed 15 µm
non-DRXed
dDRX = 3.2 µm
fDRX = 65%
Max.=5.5 DRXed
non-DRXed
15 µm
Fig. 2. Microstructural characteristics of TAZ711 alloy extruded (a–c) without and (d–f) with artificial cooling: (a, d) optical micrographs showing the recrystallized and nonrecrystallized regions, (b, e) inverse pole figure maps of the recrystallized region and corresponding ED inverse pole figures, and (c, f) XRD inverse pole figure. fDRX and dDRX represent the area fraction and average size of recrystallized grains, respectively.
alloy, with both exhibiting a partially dynamically recrystallized (DRXed) microstructure consisting of equiaxed fine DRXed grains and elongated coarse non-DRXed grains (Fig. 2a and d). As the DRX behavior is affected by alloy composition and the extrusion conditions [11], this partially DRXed structure indicates that the strain and strain rate applied during extrusion in this study is insufficient to fully induce DRX behavior in TAZ711 alloy, irrespective of the difference in deformation temperature caused by cooling. Nevertheless, even though this partial DRX structure is observed in both cases, the fraction and size of the DRXed grains differs quite significantly between the two. Specifically, the AC sample shows a decrease in the area fraction of DRXed grains (DRX fraction, fDRX) from 73% to 65% (Fig. 2a and d), as well as a decrease in the average size of the DRXed grains (DRXed grain size, dDRX) from 5.4 to 3.2 μm (Fig. 2b and e). Further observation of the microstructure of the extruded bars within the sub-surface region, which was in direct contact with the cooling water, revealed that other than a slightly enhanced cooling effect the microstructural evolution in this region was not that dissimilar to the center. This similarity between the surface and center regions can be mainly attributed to the relatively small diameter of the extruded bar (11.3 mm) and the applied feeding rate of 3 ℓ/min being sufficient to ensure the whole sample was cooled quickly. This microstructural variation with cooling is consistent with previous reports of a decrease in DRX fraction and DRXed grain size at lower extrusion temperatures [11,12,18], indicating that although the DRX fraction increases with deformation temperature due to the promotion of DRX behavior, a lower extrusion temperature leads to a reduced DRX fraction. Furthermore, the DRXed grain size is determined by the competition between DRX and grain growth during hot deformation [19], in that the rate of grain growth after DRX decreases at lower temperatures, whereas the DRXed grain size is reduced by artificial cooling. The textures of both the AC and NC samples were typical of an extruded Mg alloy, in that most grains were oriented such that their basal planes {0001} and 〈10 10〉 directions were aligned parallel to the ED (Fig. 2c and f). However, the maximum texture intensity of the AC sample (5.5) was higher than that of the NC sample (4.0), which is mainly due to a stronger texture of DRXed grains in the AC sample. The inverse pole figures in Fig. 2b and e also show that the maximum texture intensity of DRXed grains is higher in the AC sample (5.2) than the NC sample (3.1), which can be explained by the fact that the texture of the DRXed grains weakens with increasing deformation temperature due to
activation of non-basal slip systems [20]. In addition, the nonDRXed grains generally have a stronger texture than the DRXed grains [14,21], which is supported by the higher maximum texture intensities (5.5 in AC and 4.0 in NC) of the overall area (Fig. 2c and f) relative to the intensities (5.2 in AC and 3.1 in NC) acquired from only DRXed grains (Fig. 2b and e). Thus, the increase in the fraction of non-DRXed grains with a strong texture (from 27% to 35%) that is induced by cooling also clearly contributes to an intensified basal texture in the extruded AC sample. It is already known that numerous Mg2Sn particles are dynamically precipitated during the extrusion of high-Sn Mg alloys [6,7], and that the number, size and morphology of these precipitates is strongly dependent on the deformation temperature [22]. With this in mind, calculations performed using FactSage software [23] (Fig. 3a) have shown that the solubility of Sn in an aMg matrix at a homogenization temperature of 500 1C is 7 wt%, which is consistent with the complete dissolution of the Mg2Sn second phase in the homogenized TAZ711 alloy. This Sn solubility rapidly decreases with the actual temperature generated during extrusion, with any Sn content in excess of the solubility limit at a given extrusion temperature forming Mg2Sn particles during extrusion by dynamic precipitation. As shown in Fig. 3a, the solubility of Sn in the a-Mg matrix is 3.18 wt% at 370 1C, but this drops sharply to just 0.94 wt% at 260 1C (Fig. 3a). In other words, 3.82 wt% Sn is predicted to precipitate as Mg2Sn particles in the NC sample, while in the AC sample this figure is 6.06 wt%. This means that the density of Mg2Sn precipitates should increase with artificial cooling due to enhanced dynamic precipitation. This is confirmed by the SEM micrographs taken of the DRXed regions of the extruded samples (Fig. 3b and c), in which the precipitates are clearly larger in size and more prevalent in the AC sample. Furthermore, the area faction of Mg2Sn precipitates is 9.7% in the NC sample versus 14.8% in the AC sample (Table 1), with the ratio of this precipitate fraction (1.53) matching well with the ratio of precipitated Sn content between the two samples (1.59). This indicates that the drop in deformation temperature that is induced by artificial cooling effectively enhances the dynamic precipitation behavior during extrusion, subsequently resulting in more precipitates being formed. From the tensile stress–strain curves of the AC and NC samples shown in Fig. 4, it is evident that both the yield and tensile strength is increased through artificial cooling by 33 and 13 MPa, respectively (Table 1). There are several strengthening mechanisms that can occur in extruded Mg alloys, which include grain-boundary,
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fMg2Sn = 9.7%
3 µm
fMg2Sn = 14.8%
3 µm Fig. 3. (a) Variation in the solubility of Sn in an a-Mg matrix, as calculated using FactSage software. SEM micrographs showing Mg2Sn precipitates in TAZ711 alloy extruded (b) without and (c) with artificial cooling. TAC and TNC in (a) represent respectively the average billet temperatures generated during extrusion with and without artificial cooling, as obtained from Fig. 1b, and TT4 indicates the temperature of homogenization heat treatment. f Mg2 Sn in (b) and (c) represents the fraction of Mg2Sn second phases.
Table 1 Microstructural characteristics and tensile properties of TAZ711 alloy extruded with and without artificial cooling. Extruded material
Without cooling With cooling
Microstructural characteristics
Tensile properties
fDRX (%)
dDRX(μm) Imax
f Mg2 Sn
73
5.4
4.0
65
3.2
5.5
YS (MPa)
UTS (MPa)
EL (%)
9.7
226
291
6.9
14.8
259
304
5.7
(%)
fDRX and dDRX represent the fraction and average size of recrystallized grains, respectively. Imax indicates the maximum intensity of XRD inverse pole figure. f Mg2 Sn indicates the area fraction of Mg2Sn precipitates. YS, UTS, and EL represent the 0.2% offset yield strength, ultimate tensile strength, and elongation, respectively.
300
Stress (MP Pa)
250
200
150
without cooling with cooling
100
50
0 0
2
4
6
8
10
Strain (%) Fig. 4. Tensile stress–strain curves of TAZ711 alloy extruded with and without artificial cooling.
texture, dispersion/precipitate and solid-solution strengthening [1,16]; however, the improvement in strength achieved by artificial cooling in this study is considered to be the result of a combination of a refinement of the DRXed grains (from 5.4 to 3.2 μm), an enhanced texture intensity (from 4.0 to 5.5), and an increased fraction of precipitates (from 9.7% to 14.8%). The reduced DRXed
grain size increases the grain-boundary strengthening through the Hall–Petch effect by impeding the movement of dislocations and hindering the onset of plasticity. A previous study [7] has also shown that an increase in DRX fraction leads to an improved strength in extruded TAZ811 alloys due to an enhanced grainboundary strengthening effect created by the formation of fine DRXed grains ( 0.8 μm) rather than coarse non-DRXed grains. However, it should be noted that the situation is made far more complex if the extruded alloy contains relatively large DRXed grains, as this creates a conflict between texture strengthening in nonDRXed grains and grain-boundary strengthening in DRXed grains. Specifically, non-DRXed grains with a strong basal texture enhance texture strengthening under tension along the ED, even though the effects of grain-boundary strengthening are reduced by their much larger size. Consequently, the larger size of the DRXed grains in the extruded TAZ711 of this study (3.2–5.4 μm) compared to previously reported TAZ811 alloys ( 0.8 μm) indicates that the increased fraction of non-DRXed grains (from 27% to 35%) due to artificial cooling is likely to reduce the grain-boundary strengthening effect, yet improve the texture strengthening effect when tensioned along the ED. However, it is not clear at this stage which of these effects is the most predominant. It is also apparent that the increased quantity of Mg2Sn precipitates induced by artificial cooling reduces the interparticle spacing and increases the particle diameter, which in turn results in an improved precipitate strengthening effect during deformation in accordance with the Orowan equation [24]. Specifically, the stress required to move dislocations through a lattice containing precipitates is increased by greater interaction between particles and dislocations, while the stress field around particles also adds resistance to the motion of the next dislocation. The downside of this is that the increased fraction of non-DRXed grains created by artificial cooling (from 27% to 35%) causes a slight decrease in elongation (from 6.9% to 5.7%) due to the greater ease with which {10 11}–{10 12} double twins (and microcracks at these twins) are formed in larger non-DRXed grains under tension along the ED [21,25]. On the basis of these results, it is concluded that artificial water cooling during extrusion is an effective method to improve the strength of Mg–Sn based alloys, albeit at the expense of a slight loss of ductility. The relatively high temperature (350 1C) and low ram speed (1 mm/s) used for extrusion in this study led to a relatively small rise in temperature ( 20 1C) as deformation heating decreases with increasing initial temperature and decreasing extrusion speed [12]. However, a previous example of Mg–1Zn–1Mn–0.5Ce alloy extruded at 300 1C and 9 mm/s under the same indirect
S.H. Park et al. / Materials Science & Engineering A 625 (2015) 369–373
extrusion process led to a significant increase in die temperature from 300 to 380 1C [12]. Artificial cooling is therefore considered to be more effective with extrusion at low temperature and/or high speed by dramatically reducing the actual temperature of the deformation zone, but the increase in extrusion load created by this decrease in temperature (Fig. 1b) needs to be taken into account to ensure that it does not exceed the available pressure of the extrusion machinery used. Clearly then, further research is still needed to fully optimize both the extrusion conditions (e.g., temperature, speed, ratio, billet size, die angle, etc.) and cooling conditions (e.g., system design, feeding rate, media type such as water, air and gas, etc.) if the full benefits of artificial cooling with regards to the mechanical properties of extruded Mg alloys are to be realized. 4. Conclusions Artificial cooling applied immediately after extrusion has been shown to dramatically reduce the actual temperature of extruded TAZ711 alloy, which in turn reduces the size of its DRXed grains, intensifies its texture, and increases the amount of fine Mg2Sn precipitates when compared to non-cooled TAZ711. This ultimately results in a considerable increase in the yield and ultimate tensile strength of the extruded alloy through the combined effects of grain refinement, an intensified basal texture, and enhanced precipitate hardening. The elongation of the extruded alloy, however, is slightly reduced by an increased amount of large, nonDRXed grains. Nevertheless, the results obtained demonstrate that it is indeed feasible to rapidly extrude high-strength Mg alloy through a combination of a Mg–Sn based alloy developed specifically for high-speed extrusion, in conjunction with an artificial cooling system designed to enhance its strength. This therefore has the potential to meet commercial requirements for the manufacture of structural components both in terms of productivity and physical properties of the final product. Acknowledgments This work was supported by the World Premier Materials Program funded by the Korean Ministry of Knowledge Economy (No. 10037928).
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References [1] C. Bettles, M. Barnett, Advances in Wrought Magnesium Alloys: Fundamentals of Processing, Properties and Applications, Woodhead Publishing, Philadelphia, PA, USA, 2012. [2] S.H. Park, H.S. Kim, B.S. You, Korean J. Met. Mater. 51 (2013) 637–644. [3] D.H. Kang, S.S. Park, N.J. Kim, Mater. Sci. Eng. A 413 (2005) 555–560. [4] T.T. Sasaki, K. Yamamoto, T. Homma, S. Kamado, K. Hono, Scr. Mater. 59 (2008) 1111–1114. [5] F.R. Elsayed, T.T. Sasaki, T. Ohkubo, H. Takahashi, S.W. Xu, S. Kamado, K. Hono, Mater. Sci. Eng. A 588 (2013) 318–328. [6] S.S. Park, W.N. Tang, B.S. You, Mater. Lett. 64 (2011) 31–34. [7] S.H. Park, J.H. Lee, H. Yu, J. Yoon, B.S. You, Mater. Sci. Eng. A 612 (2014) 197–201. [8] S.H. Park, H.S. Kim, B.S. You, Met. Mater. Int. 20 (2014) 291–296. [9] Y. Chen, Q. Wang, J. Peng, C. Zhai, W. Ding, J. Mater. Process. Technol. 182 (2007) 281–285. [10] M. Shahzad, L. Wagner, Mater. Sci. Eng. A 506 (2009) 141–147. [11] M. Hirano, M. Yamasaki, K. Hagihara, K. Higashida, Y. Kawamura, Mater. Trans. 9 (2010) 1640–1647. [12] S.H. Park, B.S. You, R.K. Mishra, A.K. Sachdev, Mater. Sci. Eng. A 597 (2014) 296–406. [13] T. Murai, S. Matsuoka, S. Miyamoto, Y. Oki, J. Mater. Process. Technol. 141 (2003) 207–2012. [14] W.L. Cheng, H.S. Kim, B.S. You, B.H. Koo, S.S. Park, Mater. Lett. 65 (2011) 1525–1527. [15] B.C. Kim, C.H. Park, H.S. Kim, B.S. You, S.S. Park, Scr. Mater. 76 (2014) 21–24. [16] H. Yu, Y.M. Kim, B.S. You, H.S. Yu, S.H. Park, Mater. Sci. Eng. A 559 (2013) 798–807. [17] J.G. Jung, S.H. Park, B.S. You, J. Alloys Compd. http://dx.doi.org/10.1016/j. jallcom.2014.11.213. [18] H. Yu, S.H. Park, B.S. You, Mater. Sci. Eng. A 610 (2014) 445–449. [19] M. Ben-Haroush, G. Ben-Hamu, D. Eliezer, L. Wagner, Corros. Sci. 50 (2008) 1766–1778. [20] H. Koh, T. Sakai, H. Utsunomiya, S. Minamiguchi, Mater. Trans. 48 (2007) 2023–2027. [21] S.H. Park, H.S. Kim, J.H. Bae, C.D. Yim, B.S. You, Scr. Mater. 69 (2013) 250–253. [22] S.W. Xu, S. Kamado, T. Honma, Scr. Mater. 63 (2010) 293–296. [23] C.W. Bale, P. Chartrand, S.A. Degterov, G. Eriksson, K. Hack, R. Ben Mahfoud, J. Melançon, A.D. Pelton, S. Petersen, Calphad 26 (2002) 189–228. [24] E. Orowan, Symposium on Internal Stresses in Metals and Alloys, Institute of Metals, London, London, 1948. [25] M.R. Barnett, MR, Mater. Sci. Eng. A 464 (2007) 8–16.
Contents lists available at ScienceDirect
Materials Science & Engineering A journal homepage: www.elsevier.com/locate/msea
Improving the tensile strength of Mg–7Sn–1Al–1Zn alloy through artificial cooling during extrusion Sung Hyuk Park n, Ha Sik Kim, Bong Sun You Light Metal Division, Korea Institute of Materials Science (KIMS), Changwon 642-831, Republic of Korea
art ic l e i nf o
a b s t r a c t
Article history: Received 6 November 2014 Received in revised form 27 November 2014 Accepted 2 December 2014 Available online 10 December 2014
An Mg–7Sn–1Al–1Zn alloy known to have excellent extrudability and superior strength was subjected to artificial cooling during indirect extrusion by directly spraying water onto the extruded rod at the die exit. The results obtained revealed that this artificial cooling dramatically reduces the temperature of the deformation zone during extrusion, thereby creating a finer grain size, an intensified texture and a greater amount of precipitates when compared to extrusion without artificial cooling. The yield and tensile strength of the extruded alloy is also significantly improved, which is attributed to the effects of grain refinement in combination with an enhanced texture and precipitate hardening. & 2014 Elsevier B.V. All rights reserved.
Keywords: Magnesium alloy Extrusion Artificial cooling Tensile strength Precipitation
1. Introduction The automotive industry's demand for wrought magnesium (Mg) alloys has been increasing in recent years due to their mechanical superiority over comparable cast alloys, not to mention their extremely low density relative to other metallic materials. In the case of extruded Mg alloys, however, supply and application has so far been greatly limited by the low maximum extrusion speed of commercial high-strength Mg alloys such as AZ80 and ZK60 [1]; a slower manufacturing speed invariably means a higher final product cost. This has led to an increased interest in Mg–Sn based alloys due to their ability to be extruded without developing surface cracks at speeds equivalent to those used for aluminum alloys [2], which is attributed to the formation of a thermally stable Mg2Sn phase in place of low-melting temperature Mg17Al12 or MgZn2 phases [2–4]. Of all the Mg–Sn based alloys, those with a high Sn content have been the subject of particular focus by virtue of the very attractive combination of high strength and superior extrudability that they offer. Indeed, Elsayed et al. [5] have already developed an Mg–9.8Sn–3.0Al– 0.5Zn (wt%) (TAZ1031) alloy exhibiting a symmetrical tensioncompression yield strength and high ultimate tensile strength (UTS) of 358 MPa in an as-extruded state, while Sasaki et al. [4] have created an Mg–9.8Sn–1.2Zn–1.0Al (wt%) alloy with a high tensile yield strength (TYS) of 308 MPa through extrusion at a low temperature of 250 1C with a low ram speed of 0.1 mm/s.
n
Corresponding author. Tel.: þ 82 55 280 3516; fax: þ 82 55 280 3599. E-mail address: [email protected] (S.H. Park).
http://dx.doi.org/10.1016/j.msea.2014.12.011 0921-5093/& 2014 Elsevier B.V. All rights reserved.
Meanwhile, Park et al. [6] have also demonstrated an extruded Mg–7.95Sn–0.95Al–0.95Zn (wt%) (TAZ811) alloy exhibiting superior tensile and compressive strengths compared to extruded AZ31 alloy. In our previous study [7], an extraordinarily high-strength TAZ811 alloy (TYS of 390 MPa, UTS of 405 MPa) was fabricated by extruding a billet subjected to cold forging prior to extrusion. It has also been recently reported that Mg–7Sn–1Al–1Zn (wt%) (TAZ711) alloy has comparable tensile strength to commercial high-strength AZ80 alloy extruded under the same conditions [2], in addition to an excellent extrudability that is demonstrable by a maximum extrusion speed of more than 12 m/min [2,8]. It is well known that the microstructure and mechanical properties of extruded Mg alloys are strongly dependent on extrusion parameters such as billet size, die geometry, reduction ratio, initial temperature, and ram speed [8–14]. For instance, with an increase in extrusion speed, the tensile strength of the extruded alloy is gradually reduced by an increase in the final grain size. Similarly, although Mg–Sn based alloys can be rapidly extruded without cracking, such high-speed extrusion invariably induces some degree of deformation heating that reduces the strength of the extruded alloy through grain growth [8,14]. Cheng et al. [14] have also shown a decrease in the tensile strength of extruded TAZ811 alloy (from 244 to 199 MPa in TYS and from 312 to 286 MPa in UTS) with increasing extrusion exit speeds (from 2 to 10 m/min). The additional heating caused by abrupt plastic deformation and friction during high-speed extrusion should therefore be reduced in order to limit grain growth and ensure high strength in the extruded alloy. In this regard, one method that has proven effective in suppressing excessive temperature rise at high extrusion speeds is the application of artificial cooling,
370
S.H. Park et al. / Materials Science & Engineering A 625 (2015) 369–373
with Kim et al. [15] using this to achieve an increase in the tensile strength of indirectly extruded AZ31 alloy. However, there has as yet been no investigation into the effect of artificial cooling on Mg–Sn based alloys developed specifically for high speed extrusion. Moreover, although it is well known that the fine precipitates generated during extrusion play an important role in determining the strength of the final product, the influence of artificial cooling on dynamic precipitation has been overlooked by previous studies concerned only with AZ31 alloy that contains no second phases. The main objective of the present study is therefore to determine the microstructural variation that induced by artificial cooling in terms of grain size, texture and precipitate in TAZ711 that is known to have a superior strength and excellent extrudability. The effect of this variation on the tensile properties of the extruded alloy is also explored and discussed herein.
The tensile properties of the extruded samples were measured at room temperature using an Instron 4206 universal testing machine with a strain rate of 1.0 10 3 s 1. Dog-bone shaped (gage section: ∅6 mm 25 mm) specimens were used for tensile testing, the axes of which corresponded to the extrusion direction (ED). The tensile test for each sample was repeated three times to ensure repeatability and confirm the consistency of the results, but for the sake of simplicity, a representative curve for each test was used. It should also be noted that the 6 mm gage diameter of the tensile specimens represents the center of the extruded bars (∅11.3 mm), and thus their tensile properties are more closely related to the microstructure within this region than that of the sub-surface region. This means that the microstructural characteristics of the extruded bars were analyzed with a particular focus on their center region.
2. Experimental procedure
3. Results and discussion
Cast billets of TAZ711 alloy were prepared by first melting in an electric resistance furnace at 800 1C under an inert atmosphere containing a mixture of CO2 and SF6, and then pouring into a steel mold pre-heated to 210 1C. The chemical composition of the resulting billet was found through inductively couple plasma spectrometry to be 6.81Sn–1.10Al–1.07Zn (wt%). These billets were homogenized at 500 1C for 24 h to fully dissolve any macrosegregation of the solute elements and intermetallic compounds formed during solidification, and then water-quenched to obtain a supersaturated solid-solution containing no Mg2Sn phase. Using a previously described process [16], two billets (∅50 mm 150 mm) pre-heated to 350 1C were indirectly extruded at a temperature of 350 1C using a ram speed of 1 mm/s and an extrusion ratio of 20, with artificial cooling being applied to only one of these. A schematic diagram of the indirect extrusion process and associated artificial cooling system is shown in Fig. 1a. In this, cold water was fed at a rate of 3 ℓ/min through two inlets located on the surface of the stem, being then transferred through holes inside the stem and die to spray directly onto the extruded rod at the die exit. The variation in die temperature during extrusion was measured using a thermocouple installed inside the die (Fig. 1a). The microstructural and textural characteristics of the extruded samples were analyzed by optical microscopy (OM), electron backscatter diffraction (EBSD), field emission scanning electron microscopy (FE-SEM), and X-ray diffraction (XRD). The area fraction of recrystallized grains and secondary phases in each of the extruded samples was obtained by averaging the values taken from five low-magnification OM and SEM images, respectively.
Fig. 1b shows the measured variation in the extrusion load and die temperature during extrusion, the latter being considered a more accurate reflection of the actual temperature applied to the alloy during extrusion. In the case of extrusion without cooling, although the initial extrusion temperature is 350 1C, the die temperature gradually rises to 378 1C during the initial stage of extrusion due to the heat generated by plastic deformation and friction. This reduces the hot working flow stress, and subsequently leads to a decrease in the extrusion load. However, as the extrusion process stabilizes, the die temperature decreases slightly to 370 1C; a temperature that is maintained to nearly the end of extrusion. Conversely, when artificial cooling is applied during extrusion, the die temperature is rapidly reduced to 260 1C immediately after the extruded bar exits the die, this being despite the corresponding increase in extrusion load due to the greater flow stress of the billet. Furthermore, this reduced temperature remains essentially unchanged throughout whole extrusion process, which indicates that artificial cooling applied directly to the extruded bar significantly reduces its actual temperature during extrusion from 370 to 260 1C (by 110 1C). This, in turn, should result in a significant grain refinement through the suppression of grain growth. As confirmed in our previous study [17], homogenized TAZ711 alloy consists of equiaxed grains with an average size of 340 um and a very small amount of Al3Fe or Al8Mn5 inclusions; however, any coarse Mg2Sn phases formed during solidification are fully dissolved by homogenization treatment. Fig. 2 shows the microstructure and texture within the center region of the artificially cooled (AC) and naturally cooled (NC) samples of extruded TAZ711
140
600
without cooling with cooling
550
120
Die
Billet
Stem Extrudate
Die
Stem Inlet
500
Extrusion load
450
100
400
80
350
Die temperature
300
60
250
Ex xtrusion load (ton)
90o
Thermocouple Inlet
Die temperature (oC C)
Sealing element
Container
40
200 0
20
40
60 100 80 Ram distance (mm)
120
Fig. 1. (a) Schematic diagram of the indirect extrusion process and artificial cooling system used in this study. (b) Variation in die temperature and extrusion load during extrusion.
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fDRX = 73%
dDRX = 5.4 µm
ED
ED
371
Max.=4.0 non-DRXed DRXed 15 µm
non-DRXed
dDRX = 3.2 µm
fDRX = 65%
Max.=5.5 DRXed
non-DRXed
15 µm
Fig. 2. Microstructural characteristics of TAZ711 alloy extruded (a–c) without and (d–f) with artificial cooling: (a, d) optical micrographs showing the recrystallized and nonrecrystallized regions, (b, e) inverse pole figure maps of the recrystallized region and corresponding ED inverse pole figures, and (c, f) XRD inverse pole figure. fDRX and dDRX represent the area fraction and average size of recrystallized grains, respectively.
alloy, with both exhibiting a partially dynamically recrystallized (DRXed) microstructure consisting of equiaxed fine DRXed grains and elongated coarse non-DRXed grains (Fig. 2a and d). As the DRX behavior is affected by alloy composition and the extrusion conditions [11], this partially DRXed structure indicates that the strain and strain rate applied during extrusion in this study is insufficient to fully induce DRX behavior in TAZ711 alloy, irrespective of the difference in deformation temperature caused by cooling. Nevertheless, even though this partial DRX structure is observed in both cases, the fraction and size of the DRXed grains differs quite significantly between the two. Specifically, the AC sample shows a decrease in the area fraction of DRXed grains (DRX fraction, fDRX) from 73% to 65% (Fig. 2a and d), as well as a decrease in the average size of the DRXed grains (DRXed grain size, dDRX) from 5.4 to 3.2 μm (Fig. 2b and e). Further observation of the microstructure of the extruded bars within the sub-surface region, which was in direct contact with the cooling water, revealed that other than a slightly enhanced cooling effect the microstructural evolution in this region was not that dissimilar to the center. This similarity between the surface and center regions can be mainly attributed to the relatively small diameter of the extruded bar (11.3 mm) and the applied feeding rate of 3 ℓ/min being sufficient to ensure the whole sample was cooled quickly. This microstructural variation with cooling is consistent with previous reports of a decrease in DRX fraction and DRXed grain size at lower extrusion temperatures [11,12,18], indicating that although the DRX fraction increases with deformation temperature due to the promotion of DRX behavior, a lower extrusion temperature leads to a reduced DRX fraction. Furthermore, the DRXed grain size is determined by the competition between DRX and grain growth during hot deformation [19], in that the rate of grain growth after DRX decreases at lower temperatures, whereas the DRXed grain size is reduced by artificial cooling. The textures of both the AC and NC samples were typical of an extruded Mg alloy, in that most grains were oriented such that their basal planes {0001} and 〈10 10〉 directions were aligned parallel to the ED (Fig. 2c and f). However, the maximum texture intensity of the AC sample (5.5) was higher than that of the NC sample (4.0), which is mainly due to a stronger texture of DRXed grains in the AC sample. The inverse pole figures in Fig. 2b and e also show that the maximum texture intensity of DRXed grains is higher in the AC sample (5.2) than the NC sample (3.1), which can be explained by the fact that the texture of the DRXed grains weakens with increasing deformation temperature due to
activation of non-basal slip systems [20]. In addition, the nonDRXed grains generally have a stronger texture than the DRXed grains [14,21], which is supported by the higher maximum texture intensities (5.5 in AC and 4.0 in NC) of the overall area (Fig. 2c and f) relative to the intensities (5.2 in AC and 3.1 in NC) acquired from only DRXed grains (Fig. 2b and e). Thus, the increase in the fraction of non-DRXed grains with a strong texture (from 27% to 35%) that is induced by cooling also clearly contributes to an intensified basal texture in the extruded AC sample. It is already known that numerous Mg2Sn particles are dynamically precipitated during the extrusion of high-Sn Mg alloys [6,7], and that the number, size and morphology of these precipitates is strongly dependent on the deformation temperature [22]. With this in mind, calculations performed using FactSage software [23] (Fig. 3a) have shown that the solubility of Sn in an aMg matrix at a homogenization temperature of 500 1C is 7 wt%, which is consistent with the complete dissolution of the Mg2Sn second phase in the homogenized TAZ711 alloy. This Sn solubility rapidly decreases with the actual temperature generated during extrusion, with any Sn content in excess of the solubility limit at a given extrusion temperature forming Mg2Sn particles during extrusion by dynamic precipitation. As shown in Fig. 3a, the solubility of Sn in the a-Mg matrix is 3.18 wt% at 370 1C, but this drops sharply to just 0.94 wt% at 260 1C (Fig. 3a). In other words, 3.82 wt% Sn is predicted to precipitate as Mg2Sn particles in the NC sample, while in the AC sample this figure is 6.06 wt%. This means that the density of Mg2Sn precipitates should increase with artificial cooling due to enhanced dynamic precipitation. This is confirmed by the SEM micrographs taken of the DRXed regions of the extruded samples (Fig. 3b and c), in which the precipitates are clearly larger in size and more prevalent in the AC sample. Furthermore, the area faction of Mg2Sn precipitates is 9.7% in the NC sample versus 14.8% in the AC sample (Table 1), with the ratio of this precipitate fraction (1.53) matching well with the ratio of precipitated Sn content between the two samples (1.59). This indicates that the drop in deformation temperature that is induced by artificial cooling effectively enhances the dynamic precipitation behavior during extrusion, subsequently resulting in more precipitates being formed. From the tensile stress–strain curves of the AC and NC samples shown in Fig. 4, it is evident that both the yield and tensile strength is increased through artificial cooling by 33 and 13 MPa, respectively (Table 1). There are several strengthening mechanisms that can occur in extruded Mg alloys, which include grain-boundary,
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fMg2Sn = 9.7%
3 µm
fMg2Sn = 14.8%
3 µm Fig. 3. (a) Variation in the solubility of Sn in an a-Mg matrix, as calculated using FactSage software. SEM micrographs showing Mg2Sn precipitates in TAZ711 alloy extruded (b) without and (c) with artificial cooling. TAC and TNC in (a) represent respectively the average billet temperatures generated during extrusion with and without artificial cooling, as obtained from Fig. 1b, and TT4 indicates the temperature of homogenization heat treatment. f Mg2 Sn in (b) and (c) represents the fraction of Mg2Sn second phases.
Table 1 Microstructural characteristics and tensile properties of TAZ711 alloy extruded with and without artificial cooling. Extruded material
Without cooling With cooling
Microstructural characteristics
Tensile properties
fDRX (%)
dDRX(μm) Imax
f Mg2 Sn
73
5.4
4.0
65
3.2
5.5
YS (MPa)
UTS (MPa)
EL (%)
9.7
226
291
6.9
14.8
259
304
5.7
(%)
fDRX and dDRX represent the fraction and average size of recrystallized grains, respectively. Imax indicates the maximum intensity of XRD inverse pole figure. f Mg2 Sn indicates the area fraction of Mg2Sn precipitates. YS, UTS, and EL represent the 0.2% offset yield strength, ultimate tensile strength, and elongation, respectively.
300
Stress (MP Pa)
250
200
150
without cooling with cooling
100
50
0 0
2
4
6
8
10
Strain (%) Fig. 4. Tensile stress–strain curves of TAZ711 alloy extruded with and without artificial cooling.
texture, dispersion/precipitate and solid-solution strengthening [1,16]; however, the improvement in strength achieved by artificial cooling in this study is considered to be the result of a combination of a refinement of the DRXed grains (from 5.4 to 3.2 μm), an enhanced texture intensity (from 4.0 to 5.5), and an increased fraction of precipitates (from 9.7% to 14.8%). The reduced DRXed
grain size increases the grain-boundary strengthening through the Hall–Petch effect by impeding the movement of dislocations and hindering the onset of plasticity. A previous study [7] has also shown that an increase in DRX fraction leads to an improved strength in extruded TAZ811 alloys due to an enhanced grainboundary strengthening effect created by the formation of fine DRXed grains ( 0.8 μm) rather than coarse non-DRXed grains. However, it should be noted that the situation is made far more complex if the extruded alloy contains relatively large DRXed grains, as this creates a conflict between texture strengthening in nonDRXed grains and grain-boundary strengthening in DRXed grains. Specifically, non-DRXed grains with a strong basal texture enhance texture strengthening under tension along the ED, even though the effects of grain-boundary strengthening are reduced by their much larger size. Consequently, the larger size of the DRXed grains in the extruded TAZ711 of this study (3.2–5.4 μm) compared to previously reported TAZ811 alloys ( 0.8 μm) indicates that the increased fraction of non-DRXed grains (from 27% to 35%) due to artificial cooling is likely to reduce the grain-boundary strengthening effect, yet improve the texture strengthening effect when tensioned along the ED. However, it is not clear at this stage which of these effects is the most predominant. It is also apparent that the increased quantity of Mg2Sn precipitates induced by artificial cooling reduces the interparticle spacing and increases the particle diameter, which in turn results in an improved precipitate strengthening effect during deformation in accordance with the Orowan equation [24]. Specifically, the stress required to move dislocations through a lattice containing precipitates is increased by greater interaction between particles and dislocations, while the stress field around particles also adds resistance to the motion of the next dislocation. The downside of this is that the increased fraction of non-DRXed grains created by artificial cooling (from 27% to 35%) causes a slight decrease in elongation (from 6.9% to 5.7%) due to the greater ease with which {10 11}–{10 12} double twins (and microcracks at these twins) are formed in larger non-DRXed grains under tension along the ED [21,25]. On the basis of these results, it is concluded that artificial water cooling during extrusion is an effective method to improve the strength of Mg–Sn based alloys, albeit at the expense of a slight loss of ductility. The relatively high temperature (350 1C) and low ram speed (1 mm/s) used for extrusion in this study led to a relatively small rise in temperature ( 20 1C) as deformation heating decreases with increasing initial temperature and decreasing extrusion speed [12]. However, a previous example of Mg–1Zn–1Mn–0.5Ce alloy extruded at 300 1C and 9 mm/s under the same indirect
S.H. Park et al. / Materials Science & Engineering A 625 (2015) 369–373
extrusion process led to a significant increase in die temperature from 300 to 380 1C [12]. Artificial cooling is therefore considered to be more effective with extrusion at low temperature and/or high speed by dramatically reducing the actual temperature of the deformation zone, but the increase in extrusion load created by this decrease in temperature (Fig. 1b) needs to be taken into account to ensure that it does not exceed the available pressure of the extrusion machinery used. Clearly then, further research is still needed to fully optimize both the extrusion conditions (e.g., temperature, speed, ratio, billet size, die angle, etc.) and cooling conditions (e.g., system design, feeding rate, media type such as water, air and gas, etc.) if the full benefits of artificial cooling with regards to the mechanical properties of extruded Mg alloys are to be realized. 4. Conclusions Artificial cooling applied immediately after extrusion has been shown to dramatically reduce the actual temperature of extruded TAZ711 alloy, which in turn reduces the size of its DRXed grains, intensifies its texture, and increases the amount of fine Mg2Sn precipitates when compared to non-cooled TAZ711. This ultimately results in a considerable increase in the yield and ultimate tensile strength of the extruded alloy through the combined effects of grain refinement, an intensified basal texture, and enhanced precipitate hardening. The elongation of the extruded alloy, however, is slightly reduced by an increased amount of large, nonDRXed grains. Nevertheless, the results obtained demonstrate that it is indeed feasible to rapidly extrude high-strength Mg alloy through a combination of a Mg–Sn based alloy developed specifically for high-speed extrusion, in conjunction with an artificial cooling system designed to enhance its strength. This therefore has the potential to meet commercial requirements for the manufacture of structural components both in terms of productivity and physical properties of the final product. Acknowledgments This work was supported by the World Premier Materials Program funded by the Korean Ministry of Knowledge Economy (No. 10037928).
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