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Effect of normalizing temperature on microstructure and mechanical properties of a Nb-V microalloyed large forging steel
Author’s Accepted Manuscript Effect of normalizing temperature on microstructure and mechanical properties of a NbV microalloyed large forging steel Xin-li Wen, Zhen Mei, Bo Jiang, Li-chong Zhang, Ya-zheng Liu www.elsevier.com/locate/msea
PII: DOI: Reference:
S0921-5093(16)30717-1 http://dx.doi.org/10.1016/j.msea.2016.06.059 MSA33805
To appear in: Materials Science & Engineering A Received date: 26 May 2016 Revised date: 19 June 2016 Accepted date: 20 June 2016 Cite this article as: Xin-li Wen, Zhen Mei, Bo Jiang, Li-chong Zhang and Yazheng Liu, Effect of normalizing temperature on microstructure and mechanical properties of a Nb-V microalloyed large forging steel, Materials Science & Engineering A, http://dx.doi.org/10.1016/j.msea.2016.06.059 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Effect of normalizing temperature on microstructure and mechanical properties of a Nb-V microalloyed large forging steel Xin-li Wen, Zhen Mei, Bo Jiang, Li-chong Zhang, Ya-zheng Liu* School of Material Science and Engineering, University of Science and Technology Beijing, Beijing 100083, China * Corresponding author. [email protected] Abstract The microstructure of a microalloyed large forging steel with different normalizing temperatures ranging from 820 ºC to 940 ºC were characterized. The evolution of austenite formation was determined in a range of heating temperature from 730 ºC to 940 ºC. The mechanical properties were evaluated by tensile test and Charpy V-notch impact test. The relationship between the microstructure and the properties was discussed. The results indicated that the microstructure composed of fine-grained layers (FGL) and coarse-grained layers (CGL) was obtained at 820 ºC. The finest and most homogeneous microstructure and optimal comprehensive mechanical properties were obtained at the normalizing temperature 880 ºC. There was a Hall-Petch relationship between the yield strength and the average grain size, and a linear relationship between the impact energy(KV2) and the reciprocal of the square root of the grain size (D-1/2). Both the strength and toughness of the steel can be attributed to grain refinement. Keywords: Normalizing temperature, Microstructure, Mechanical properties, Nb-V microalloyed, Large forging steel, grain refinement.
1. Introduction Microalloyed large forging steel has been widely used for engineering components which require high strength, good toughness and large size. Large forging steel is usually made from large casting ingots. The as-forged microstructure of the steel is generally composed of coarse ferrite and banded pearlite which consequently lead to limited impact toughness. Due to the characteristic of forging process and the large volume of forging, conventional technology such as thermo-mechanical treatment or accelerated cooling is infeasible for grain refinement in order to improve toughness. However, it is likely to increase toughness levels of microalloyed large forging steel through grain refinement by normalizing. Zhao et al [1] studied the effect of W addition and normalizing conditions on microstructure and mechanical properties of microalloyed forging steels, four kinds of microalloyed forging steels were produced by varying W additions(0, 0.5, 1 and 2 wt%), heat treatment was carried out at temperatures ranging from 840 ºC to 950 ºC followed by air and furnace cooling, the results showed that the microstructure and mechanical properties of the microalloyed forging steels were closely related to the W content, normalizing temperature and cooling method after normalizing. Zhao et al [2] studied the effect of hot forging, normalizing temperature (840 ºC-950 ºC) and cooling method (air and furnace cooling) after normalizing on the toughness and tensile properties of a microalloyed cast steel, the results showed that remarkable improvement in toughness and tensile properties can be obtained by hot forging, proper normalizing temperature and air cooling after normalizing. Zhao et al [3] studied the effect of normalizing temperature (950 ºC-1200 ºC) and cooling method (furnace, air and water cooling) after normalizing on the toughness and tensile properties of a low-carbon microalloyed cast steel, the results showed that heat treatment at 1100 ºC for 2 hours followed by furnace cooling leaded
to the best combination of excellent Charpy impact and tensile properties. The above research focused on the low carbon Nb-Ti microalloyed steel, there was no many studies on Nb-V microalloyed steel in which the V content can be as much as 0.1 wt%. Besides, the sizes of samples used by Zhao et al [1-3] were 11 ×11 × 60 mm3 or 11 ×11 × 110 mm3, the air cooling rate was above 3ºC/s. There were rare studies on large section forging steel with the diameter larger than 250 mm and the air cooling rate below 0.05ºC/s. Austenite formation in low carbon steels has been studied extensively in the literature starting from different microstructures [4-6]. Previous work has shown that in ferritic-pearlitic microstructures the formation of austenite was described as taking place in three main successive steps: (1) nucleation of austenite in pearlite colonies, ferrite-pearlite grain boundaries or ferrite-ferrite grain boundaries, (2) rapid growth of austenite consuming pearlite, (3) slower growth of austenite consuming ferrite [7,8]. Based on the above theory, as for ferrite-pearlite banded microstructure in large forging steel, austenization in ferrite bands and pearlite bands are asynchronous. Nonetheless, rare literature has studied this phenomenon. What’s more, there were no many studies on the effect of intercritical normalizing on microstructure and mechanical properties of Nb-V microalloyed large forging steel. For large forging steel, it is impracticable to enhance cooling rate in case of thermal stress-cracking. The cooling method followed normalizing is usually air cooling, hence the austenization temperature is the decisive normalizing parameter for microstructure and properties. The study of this paper aims at investigating the effect of normalizing temperature on microstructure and properties of a Nb-V microalloyed large forging steel. The evolutions of austenite, microstructure and precipitations of the tested steel were characterized. The relations between the microstructure and properties were discussed. The tested steel in this work with the diameter of Φ290 mm and the V content up to 0.095 wt% has never been investigated. 2. Experimental material and procedures The steel used in this work is a commercial HSLA steel. The chemical composition is listed in Table 1. The round bar specimens with a length of 200mm and diameter 290 mm for normalizing is shown in Fig. 1. They were cut from a Φ290 mm round forging. In order to study the characteristic of microstructure at the corresponding normalizing temperature, cubic samples for quenching were wire-cut from the 1/2 radius of the Φ290 mm round forging. The size of the quenching samples is 10 ×10 × 12 mm3. Both the normalizing and quenching process was conducted in a 45 kw box resistor-stove, the schedules of the process are given in Fig. 2. The normalizing specimens were reheated at 820 ºC, 850 ºC, 880 ºC, 910 ºC and 940 ºC with soaking for 2 hours, respectively, and then were cooled by air with about a 0.03 ºC/s cooling rate. In order to study the evolution of austenite, samples for interrupted heating by quenching were respectively reheated at 730 ºC, 760 ºC, 790 ºC, 820 ºC, 850 ºC, 880 ºC, 910 ºC and 940 ºC holding for 2 hours. After the normalizing process, blanks for metallographic observation and mechanical property test were wire-cut from the 1/2 radius of the normalizing samples along longitudinal axial direction as shown Fig. 1. Metallographic observation direction for all test samples was parallel to the longitudinal section of the Φ290 mm round forging. The microstructure of the samples was etched by a 4% nital solution. The size and area fraction of ferrite and pearlite constituent were measured by software Image-Pro Plus. For each specimen, at least 5 fields of view containing at least 400 grains were measured. Thin foils for Transmission Electron Microscopy (TEM) were prepared using the twin-jet method and observed in a JEM-2100 transmission electron microscope.
The average size and fraction of precipitate particles were statistically measured by averaging 5 fields of view containing at least 300 particles from the images of TEM. In order to study the orientation characteristic of acicular pearlite, electron back scattered diffraction examinations were performed on a field emission gun scanning electron microscope. The blanks for tensile test were machined into standard tensile test specimens of 10 mm in gage diameter and 50 mm in gage length. Tensile tests based on standard of ISO 6892-1: 2009 were carried out on a WDW-200D tensile testing machine at room temperature with a cross-head speed of 0.25 mm min-1 [9]. The yield strength was determined by the 0.2% offset flow stress. All results were repeated for three times and the average values were taken to describe the tensile properties of the test steel. The Charpy V-notched specimens with cross section of 10 × 10 mm2, length of 55 mm, notch angle of 45° and notch depth of 2 mm were employed to study the -40 ºC impact fracture toughness on a ZBC2452-B impact testing machine according to ISO 148-1: 2006. Table 1 Chemical compositions of the tested steel (wt%). C
Si
Mn
P
S
Nb
V
Ti
0.15
0.27
1.45
0.01
0.004
0.045
0.095
0.01
200 mm
Mechanical sample
Metallographic sample 290 mm
Fig. 1 The normalizing sample and sampling method T=820 ºC, 850 ºC, 880 ºC, 910 ºC, 940 ºC
Temperature (ºC)
t=2h
Air cooling
10 ºC /min
about 0.03 ºC /s
Quenching Time (s)
Fig. 2 Normalizing and quenching process 3. Results and discussion 3.1. The evolution of austenite formation The study of the austenite formation was carried out using a DIL805A high resolution dilatometer. Cylindrical samples of 4 mm diameter and 10 mm length were used for the experiments. As shown in Fig. 3, the relations between heating temperature(T) and expansion amount(ΔL) were analyzed to determine Ac1 and Ac3 at a constant heating rate of 10 ºC/min. Since
some investigations have experimentally shown that a separation can be made between the pearlite to austenite and the ferrite to austenite transformation [10-12], an attempt was made to determine the temperature (Acθ) at which this occurred. The determination of Acθ can be only attempted if the first contraction is perceived. It is less evident to see from the dilatometric curve, but easier to determine from the first derivative (dΔL/dT) as shown in Fig. 3. In previous papers studied similar steels, the authors showed that this first contraction was related to the dissolution of pearlite[13, 14]. As shown in Fig. 4(b) and 6(b), microstructure morphologies support the fact that the first contraction observed in the dilatometric curve is due to pearlite dissolution.
Fig. 3 Dilatometric curve and the first derivative According to the results of dilatometer tests, the Ac1 and Ac3 of the tested steel were determined as 725 ºC and 861 ºC, respectively. Fig. 4 shows microstructures obtained after interrupted heating by quenching at 730-940 ºC. It can be seen that the specimen heated at 730 ºC mainly consists of white band and black band microstructure. With increasing temperature from 730 ºC to 850 ºC, the black band microstructure gets wider and wider, the white band gets narrower and narrower accordingly. There is no obvious banded microstructure in samples heated at 880, 910, and 940 ºC. (a)
(b)
(c)
(e)
(f)
Initial pearlite
Initial ferrite (d)
(g)
(h)
Fig. 4 Optical micrographs of quenched microstructure after heating at different temperatures (a) 730 ºC, (b) 760 ºC, (c) 790 ºC, (d) 820 ºC, (e) 850 ºC, (f) 880 ºC, (g) 910 ºC, (h) 940 ºC Fig. 5 shows the SEM micrographs of the quenched microstructure. Fig. 5(a)-Fig. 5(e) corresponds to the micrographs of the white band microstructure in Fig. 4(a)-Fig. 4(e). Fig. 5(f)-Fig. 5(h) are typical micrographs corresponding to Fig. 4(f)-Fig. 4(h). It can be seen from Fig. 5(a) that the microstructure mainly consists of ferrite as well as a small amount of martensite which are distributed at the ferrite boundaries. Formation of martensite in ferrite band indicates that actual austenization has started at 730 ºC. With increasing temperature to 760 ºC, net-like martensite can be seen clearly along the ferrite boundaries, which shows that ferrite boundaries is the preferential site for austenite nucleation and growth. With further increasing of temperature from 760 ºC to 790 ºC, the net-like martensite grows into the ferrite. Accordingly, the amount of ferrite is reduced remarkably. Fig. 5(d) and Fig. 5(e) show that the microstructure is composed of martensite and a small amount of ferrite after quenching at 820 ºC and 850 ºC, respectively. The acicular microstructure adjacent to the ferrite is turned out to be martensite as shown in Fig. 5(i). The microstructure completely consists of martensite when the heating temperature is 880 ºC, 910 ºC, and 940 ºC, indicating that the actual austenization of ferrite band has been finished at about 880 ºC. The formation of acicular microstructures (finger-type austenite) during austenization was also observed by many researchers [15-18]. It is quite possible that the formation of the fingers coincide with the position of former cementite plates that were perpendicular to the grain boundaries. (a)
(b)
ferrite
martensite
(c)
ferrite martensite (d)
martensite
ferrite
martensite
(e)
(f)
ferrite ferrite
(g)
martensite (h)
martensite (i)
Fig. 5 SEM micrographs of quenched microstructure after heating at different temperatures (a) 730 ºC, (b) 760 ºC, (c) 790 ºC, (d) 820 ºC, (e) 850 ºC, (f) 880 ºC, (g) 910 ºC, (h) 940 ºC, (i) micrographs of the acicular microstructure(TEM) Fig. 6 shows the SEM microstructures of the quenched microstructure. Fig. 6(a)-Fig. 6(e) correspond to the micrographs of the black band microstructure in Fig. 4(a)-Fig. 4(e). Fig. 6(f)-Fig. 7(h) are typical micrographs corresponding to Fig. 4(f)-Fig. 4(h). It can be seen from Fig. 6(a) that the microstructure consists of pearlite and ferrite, as well as a small amount of martensite which is distributed at the ferrite-pearlite boundaries. With increasing temperature to 760 ºC, blocky martensite can be seen clearly around ferrite, and pearlite has also been dissolved simultaneously. Fig. 6(c) and Fig. 6(d) show that the microstructure is composed of martensite and a small amount of ferrite after quenching at 790 ºC or 820 ºC. The microstructure completely consists of martensite when the heating temperature is 850, 880, 910, or 940 ºC, indicating that the actual austenization of the pearlite bands has been finished at 850 ºC. (a)
(b)
(c) ferrite ferrite
martensite pearlite ferrite
martensite
(d)
(e)
(f)
ferrite
(g)
(h)
Fig. 6 SEM micrographs of quenched microstructures after heating at different temperatures (a) 730 ºC, (b) 760 ºC, (c) 790 ºC, (d) 820 ºC, (e) 850 ºC, (f) 880 ºC, (g) 910 ºC, (h) 940 ºC Under the given conditions in this paper, the formation of austenite in initial ferrite and pearlite was asynchronous. The nucleation and growth of austenite in initial pearlite was earlier and faster than that in ferrite. The microstructure heredity of austenite on ferrite and pearlite will be shown below. 3.2. Microstructure of normalized samples Fig. 7 shows the optical micrographs of the samples before and after normalizing. As can be seen in Fig. 7(a), as-forged microstructure consists of alternate bands of proeutectoid ferrite and pearlite. EPMA was used to investigate the chemical characteristic across the banded microstructure in Fig. 7(a). Figs. 8(a) and 8(b) show manganese and silicon profiles, respectively. Excellent correlation between microstructural banding and microchemical banding is apparent. In other words, solute lean regions and solute rich regions are consistently associated with regions of proeutectoid ferrite and pearlite, respectively. The results are in full agreement with the findings of
Thompson and Howell [19]. They showed that microchemical banding of manganese was the cause of the microstructural banding observed in a similar steel. (a)
(b)
FGL
(c)
CGL
(d)
(e)
(f)
Fig. 7 The optical micrographs for the specimens before and after normalizing (a) as-forged, (b) 820 ºC, (c) 850 ºC, (d) 880 ºC, (e) 910 ºC, (f) 940 ºC
Fig. 8 Concentration profiles from as-forged sample (a) manganese profile, (b) silicon profile In an attempt to determine the reliability of the data presented in Fig. 8(a), the maximum and minimum concentrations of manganese were estimated using the Scheil equation [20], which was developed for the limiting case of no diffusion in the solid, but complete mixing in the liquid. Cs = kC0(1- fs)k-1 (1) where Cs is the concentration of solute in the solid at a given fraction solidified fs, C0 is the concentration of solute in the alloy, and k is the equilibrium partition ratio, defined as k = Cs/C1 (2) where C1 is the concentration of solute in the liquid in equilibrium with the solid of concentration Cs. To simplify the analysis, k is assumed to be constant throughout the solidification range. For manganese segregation k=0.71 and C0 is 1.45% Mn [21]. Thus, Cs=1.06%Mn for fs=0.1, and Cs=2.00%Mn for fs =0.9. These values are in reasonably good agreement with the maximum and minimum values of manganese concentration shown in Fig. 8(a). Fig. 7(b) and Fig. 9(a) show the inhomogeneous banded microstructure which consists of fine grain bands (FGB) and coarse grain bands (CGB). The FGB is composed of relatively large size of ferrite (initial ferrite) and acicular black microstructure. The CGB consists of relatively fine ferrite (F) and pearlite. The acicular microstructure as shown in Fig. 9(b) is identified as lamellar pearlite by using SEM. Hence, it is called acicular pearlite (AP) in this paper. Color coded inverse
pole figure (IPF) maps of ferrite in CGB is given in Fig. 9(c), which indicates that the adjacent acicular pearlite possesses the same crystal orientation. (a)
FGB
(c)
(b)
AP AP F CGB
Fig. 9 The microstructures obtained with the normalizing temperature 820 ºC (a) OM , (b)SEM, (c) IPF map From the above, complete austenization was finished at about 880 ºC, so there is coarse initial ferrite in the samples normalized at 820 ºC and 850 ºC. When the temperature is 880 ºC, the microstructure is the finest. With increasing temperature from 880 ºC to 940 ºC, abnormal growth of some pearlite nodules occurs, as shown in Fig. 7(e) and (f). The average size of the pearlite nodules in the specimens normalized at 910 ºC and 940 ºC are about 38 and 44 μm in diameter, respectively. These values are very similar to the largest austenite grain sizes in the specimens. As shown in Fig. 10(a), the largest austenite grain size is about 30 μm in diameter in specimen reaustenitized for 2 hours at 880 ºC. The number density is very low. Conversely, as shown in Fig. 10(b) and (c), numerous large austenite grains with the diameter ranging from 40 to 60 μm were observed in the specimens reaustenitized for 2 hours at 910 ºC and 940 ºC. (a)
(b)
(c)
Fig. 10 The micrographs for the specimens reaustenitized at different temperatures (a) 880 ºC, (b) 910 ºC, (c) 940 ºC The similarity in sizes and number densities for the large pearlite nodules and large austenite grains strongly suggests that the large pearlite nodules form within large austenite grains. If it is assumed that air cooling (about 0.03 ºC/s) produces microstructure that are relatively close to equilibrium, and it is further assumed that the pearlite nodules form from a single austenite grain, then the size of the nodules can be estimated from the austenite grain size and the volume fraction of pearlite. If, for simplicity, the austenite grains and the pearlite nodules are assumed to be spherical, then for a pearlite volume fraction of 0.39 and an austenite grain diameter of 60 μm, the diameter of the pearlite nodules would be about 44 μm. Fig. 7(f) and Fig. 10(c) show that large pearlite nodules are about 30-50 μm in diameter, consistent with formation within the largest austenite grains. Hence, it can be concluded that large, irregular pearlite nodules form in abnormally large austenite grains [22]. 3.3. Precipitation Fig. 11 shows the bright field TEM micrographs and EDS analysis of the representative precipitations observed in the as-forged specimen. The observed precipitations can be classified
into three different size ranges: 1-3 μm (cuboidal precipitations), 50-150 nm (spherical precipitations) and 5-15 nm (fine precipitations). The cuboidal precipitations are titanium-rich and niobium-contained (Ti, Nb)C (Fig. 12a and d). The spherical precipitations are niobium-rich and titanium-contained (Nb, Ti)C (Fig. 12b and e) and present a chain-like distribution. The fine precipitations are vanadium-rich and niobium-contained (V, Nb)C (Fig. 12c and f). The majority of the precipitations are duplex-type carbides. Statistics shows that there is no obvious difference of the precipitations in size and distribution between the specimens before and after normalizing. Fig. 12 shows the bright field TEM micrographs of the representative precipitations observed in the normalized specimens.
(b)
(a)
(c)
(Nb, Ti)C (V, Nb)C (Ti, Nb)C
(d)
(e)
(f)
Fig. 11 Three typical TEM micrographs and EDS analysis of the precipitations (a and d)cuboidal precipitation, (b and e)spherical precipitations, (c and f)fine precipitations (a)
(b)
(c)
(Nb, V)C
(V, Nb)C
(Ti, Nb)C
Fig. 12 Three typical TEM micrographs of the precipitations in the specimens normalized at 850 ºC (a)cuboidal precipitation, (b)spherical precipitations, (c)fine precipitations 3.3. Mechanical properties Performance requirements for the tested steel are listed here: Rp0.2≥265 MPa, Rm=450-600 MPa, A≥17%, KV2≥27 J(-40 ºC). Mechanical properties of the tested steels are ineligible before normalizing as listed: Rp0.2=326 MPa, Rm=520 MPa, A=25%, KV2=15J. The heat treated specimens exhibited superior mechanical properties over the original untreated specimen in terms of higher tensile strength, elongation and impact energy as summarized in Table 2.
Table 2 Mechanical properties of the specimens varied with normalizing temperature. T(ºC)
Rp0.2(MPa)
Rm(MPa)
A(%)
-40 ºC KV2 (J)
820
340
540
27
34
850
360
569
32
95
880
372
577
33
123
910
352
567
30
60
940
348
558
29
53
4. Discussion 4.1. Effect of microstructure refinement on yield strength It can be seen in Table 2 that mechanical properties are dramatically influenced by the normalizing temperature. With the increase of normalizing temperature, yield strength (Rp0.2), tensile strength (Rm) and elongation (A) first rise and then drop. The optimum tensile properties are obtained at the normalizing temperature 880 ºC. It is well known that the yield strength for low carbon microalloyed steels is influenced by different strengthening mechanisms, which include solid solution strengthening, dislocation strengthening, fine-grain strengthening and precipitation strengthening [23]. These factors contribute to the yield strength of steel in varying degrees, depending on the microstructure generated by the specific processing technology. Compared with as-forged samples, the yield strength of normalized samples increases 20-83 MPa, This phenomenon is considered to be explained by combination effects of several strengthening mechanisms. Precipitation strengthening can be calculated by Ashby-Orowan equation [24]:
p
0.538Gbf 1/ 2 x ln x 2b
(2)
Where σp is the increase in yield strength (MPa), G is the shear modulus (81600 MPa for Fe), b is the Burgers vector (0.248 nm), f is the volume fraction of particles, and x is the real (spatial) diameter of the precipitates (nm). Using the value of the area fraction and size distribution of precipitates, the precipitation strengthening of samples can be evaluated as Table 3. The subscript 1 represents the fine precipitations and the subscript 2 represents the spherical precipitations. The precipitation strengthening of (Ti, Nb)C can be ignored based on the fact that the large size is not very useful in increasing the strength. Compared with the as-forged samples, the calculated yield strength increments of the normalized samples due to precipitation strengthening are estimated about -7 MPa (820 ºC), -8 MPa (850 ºC), 1 MPa (880 ºC), -2 MPa (910 ºC) and 5 MPa (940 ºC), respectively, as shown in Fig. 13. The actual yield strength increments according to Table 2 are 14 MPa (820 ºC), 34 MPa (850 ºC), 46 MPa (880 ºC), 26 MPa (910 ºC) and 22 MPa (940 ºC), respectively. Comparison of the calculated and actual yield strength increments is shown in Fig. 13, it can be seen that there is no strong correlations between the yield strength change and precipitation strengthening. So the yield strength change cannot be attributed to precipitation strengthening.
Table 3 Average precipitate size, precipitation volume fraction and corresponding precipitation strengthening value of the test specimens. T(ºC)
x1(nm)
f1
σp1(MPa)
x2(nm)
f2
σp2(MPa)
σp1+σp2(MPa)
820
10.6
0.00092
95
112.2
0.00032
9
105
850
11.2
0.00095
93
110.5
0.00037
10
104
880
9.5
0.0009
102
100.6
0.00039
11
113
910
10.5
0.001
100
97.8
0.0003
10
110
940
9.2
0.00094
106
105.4
0.0004
11
117
Fig. 13 Yield strength increment by precipitation Yield strength(Rp0.2) and average grain size(D) of the specimens normalized at different temperatures are shown in Fig. 14. With the increase of normalizing temperature, yield strength first rises and then drops, but average grain size first drops and then rises, indicating that there is a negative correlation between them. The quantitative relationship between the yield strength and grain size has been known as the classic Hall–Petch equation[25,26]： σy=KyD-1/2+σ0 (2) where σy is the yield strength (MPa), σ0 is the frictional stress (MPa), Ky is the Hall–Petch slope, D is the average grain size. According to Eq. (2), fine-grained microstructure has a pronounced effect on the increment of yield strength. In another supporting work, similar conclusion was obtained by Pickering[27] with an empirical equation for ferritic-pearlitic steels under 0.25% C. The method of linear regression is introduced to analyze the relationship between Rp0.2 and D-1/2, the results are shown in Fig. 15. There is a linear relationship between Rp0.2 and D-1/2, so the yield strength change can be mainly attributed to fine-grain strengthening. Similar phenomena can be seen in several papers[28-30].
Fig. 14 Yield strength and grain size of the specimens normalized at different temperatures
Fig. 15 Relationship between yield strength and grain size 4.2. Effect of microstructure refinement on the impact toughness As shown in Table 2, with the increase of normalizing temperature, -40ºC impact energy (KV2) first rises and then drops, the optimum toughness can be obtained at the normalizing temperature 880 ºC. Toughness in steels is controlled by different microstructural constituents that can be considered as material flaws. Some of them such as inclusions are intrinsic while others depend on processing conditions, i.e., boundaries, grain boundary carbides or hard second phases. It can be seen from Fig. 16 that a linear relationship exists between the -40ºC impact energy(KV2) and the reciprocal of the square root of the grain size (D-1/2), indicating the dependence of the impact energy on the average grain size. The impact energy increases with decreasing average grain size. The impact energy is increased nearly by 3 times from 34 J to 123 J when the grain size was refined from 24 μm to 13 μm, indicating that grain refinement contributes largely to the toughness of the steel. The effect of grain refinement on the toughness of the steel can be explained by the generalized Griffith’s equation given below[31-33].
f
2 E (1 v 2 ) D
(3)
where, σf is the cleavage fracture stress, E is the Young’s Modulus, v is the poisson’s ratio, γ is the corresponding effective surface energies for crack propagation across interfaces, D is the grain size.
Fig. 16 Relationship between impact energy and grain size The fractured surfaces of the steel mainly consist of cleavage facets. The representative fracture morphologies of the tested steel normalized at 820 ºC, 880 ºC and 940 ºC are shown in Fig. 17. The arithmetic mean of the observed facet size was determined by linear intercept method on SEM images, the results are given in Table 4. For the specimens with the grain sizes ranging from 24 μm to 13 μm, the arithmetic mean of the cleavage facet size is decreased from
30 μm to 11μm. It implies that the dominant microstructural feature in the toughness of the steel is the grain size. (a)
(b)
(c)
Fig. 17 Fracture surfaces of the impact specimens normalized at (a)820 ºC; (b)880 ºC; (c)940 ºC Table 4 Average grain size and cleavage facet size of the tested specimens. T(ºC)
Grain size(μm)
Cleavage facet size(μm)
820
24
30
850
16
15
880
13
11
910
18
21
940
20
25
Further evidence of the direct relationship between the grain size and the facet size is shown in Fig. 18, which is an IPF map given by EBSD. The figure shows the crack propagation path in the vicinity of the fracture of the broken impact specimen. The different color in Fig. 18 represents the different grain orientations. The black heavy line represents the high-angle grain boundaries(≥15°), the black light line represents the low-angle grain boundaries(＜15°). Fig. 18 shows that the direction of the cleavage crack is largely changed at the high-angle grain boundaries. A local cleavage cracking event across a grain could lead to an instant load drop. However, when the crack attempts to run from one grain to another, the crystallographic orientation and microstructure may change so that more work had to be done to cross the grain boundary. As a result, the local cleavage crack might be arrested [29]. From the above results, it can be concluded that the unit crack path for cleavage fracture is identified as the grain. The grain refinement is very effective in improving the resistance to the cleavage fracture.
Crack
Fig. 18 The crack propagation path of the fractured impact specimen 5. Conclusion The effect of normalizing temperature on microstructure and mechanical properties of a Nb-V-microalloyed large forging steel has been studied in this work. The conclusions are as
follows: (1) When the normalizing temperature was in γ+α dual phase region, the original pearlite band of the tested steel can achieve complete austenization and transformed into relatively fine grain band during air cooling followed, the original ferrite band retained partly and transformed into coarse grain band subsequently. When the normalizing temperature was in γ region, relatively homogeneous ferrite and pearlite grains can be obtained. With the increase of normalizing temperature, both the ferrite and pearlite grain size increase gradually. (2) With the increase of normalizing temperature, strength and toughness of the steel first rise and then drop, the excellent combination of mechanical properties were obtained at the normalizing temperature 880 ºC. The improvement of yield strength and impact energy can be mainly attributed to grain refinement. (3) There is a Hall-Petch relationship between the yield strength and the grain size, and a linear relationship between the impact energy (KV2) and the reciprocal of the square root of the grain size (D-1/2). The strength and impact toughness of the steel can be improved with the refinement in grain size. The unit crack path for cleavage fracture has been identified as the grain. Grain refinement is very effective in improving the resistance to the cleavage fracture. Acknowledgements The authors appreciate the financial support by the National High Technology Research and Development Program (“863”Program) of China (No.2012AA03A503). References [1] Zhao J, Jiang Z, Lee C S. Effects of tungsten addition and heat treatment conditions on microstructure and mechanical properties of microalloyed forging steels[J]. Materials Science and Engineering: A, 2013, 562: 144-151. [2] Zhao J, Jiang Z, Lee C S. Enhancing impact fracture toughness and tensile properties of a microalloyed cast steel by hot forging and post-forging heat treatment processes[J]. Materials & Design, 2013, 47: 227-233. [3] Zhao J, Lee J H, Kim Y W, et al. Enhancing mechanical properties of a low-carbon microalloyed cast steel by controlled heat treatment[J]. Materials Science and Engineering: A, 2013, 559: 427-435.) [4] Li H, Gai K, He L, et al. Non-isothermal phase-transformation kinetics model for evaluating the austenization of 55CrMo steel based on Johnson–Mehl–Avrami equation[J]. Materials & Design, 2016, 92: 731-741. [5] Chang W, Cao W, Yun H A N, et al. Influences of Austenization Temperature and Annealing Time on Duplex Ultrafine Microstructure and Mechanical Properties of Medium Mn Steel[J]. Journal of Iron and Steel Research, International, 2015, 22(1): 42-47. [6] Li S, Kang Y, Zhu G, et al. Austenite formation during intercritical annealing in C-Mn cold-rolled dual phase steel[J]. Journal of Central South University, 2015, 22: 1203-1211. [7] Savran V I, Offerman S E, Sietsma J. Austenite nucleation and growth observed on the level of individual grains by three-dimensional X-ray diffraction microscopy[J]. Metallurgical and Materials Transactions A, 2010, 41(3): 583-591. [8] Souza M M, Guimaraes J R C, Chawla K K. Intercritical austenitization of two Fe-Mn-C steels[J]. Metallurgical Transactions A, 1982, 13(4): 575-579. [9] ISO 6892-1, Metallic Materials-TensileTesting-Part 1: Method of Testat Room Temperature, 2009, ISO E N. 6892-1, Metallic materials–Tensile testing Part 1: Method of test at ambient
temperature[J]. British Standard Institution, 2009. [10] Jayaswal S K, Gupta S P. Kinetics of ferrite to austenite transformation in a high strength low alloy steel containing Ti and V[J]. Zeitschrift für Metallkunde, 1992, 83(11): 809-819. [11] Speich G R, Szirmae A, Richards M J. Formation of austenite from ferrite and ferrite-carbide aggregates[J]. TRANS MET SOC AIME, 1969, 245(5): 1063-1074. [12] Speich G R, Demarest V A, Miller R L. Formation of austenite during intercritical annealing of dual-phase steels[J]. Metallurgical Transactions A, 1981, 12(8): 1419-1428. [13] Caballero F G, Capdevila C, Andrés C G D. Modelling of kinetics of austenite formation in steels with different initial microstructures[J]. ISIJ international, 2001, 41(10): 1093-1102. [14] Martín S D, Cock T, García-Junceda A, et al. Effect of heating rate on reaustenitisation of low carbon niobium microalloyed steel[J]. Materials Science and Technology, 2008, 24(3): 266-272. [15] Jayaswal S K, Gupta S P. Kinetics of ferrite to austenite transformation in a high strength low alloy steel containing Ti and V[J]. Zeitschrift für Metallkunde, 1992, 83(11): 809-819. [16] Souza M M, Guimaraes J R C, Chawla K K. Intercritical austenitization of two Fe-Mn-C steels[J]. Metallurgical Transactions A, 1982, 13(4): 575-579. [17] Savran V I, Van Leeuwen Y, Hanlon D N, et al. Microstructural features of austenite formation in C35 and C45 alloys[J]. Metallurgical and Materials Transactions A, 2007, 38(5): 946-955. [18] Martín S D, Cock T, García-Junceda A, et al. Effect of heating rate on reaustenitisation of low carbon niobium microalloyed steel[J]. Materials Science and Technology, 2008, 24(3): 266-272. [19] Thompson S W, Howell P R. Factors influencing ferrite/pearlite banding and origin of large pearlite nodules in a hypoeutectoid plate steel[J]. Materials Science and Technology, 1992, 8(9): 777-784. [20] Scheil E. Unbroken series of solid solutions in the binary systems of the elements[J]. Z. Metallkd, 1942, 34: 242-246. [21] Thompson S W, Howell P R. Factors influencing ferrite/pearlite banding and origin of large pearlite nodules in a hypoeutectoid plate steel[J]. Materials Science and Technology, 1992, 8(9): 777-784. [22] Majka T F, Matlock D K, Krauss G. Development of microstructural banding in low-alloy steel with simulated Mn segregation[J]. Metallurgical and Materials Transactions A, 2002, 33(6): 1627-1637. [23] Han Y, Shi J, Xu L, et al. TiC precipitation induced effect on microstructure and mechanical properties in low carbon medium manganese steel[J]. Materials Science and Engineering: A, 2011, 530: 643-651. [24] Gladman T. Precipitation hardening in metals[J]. Materials science and technology, 1999, 15(1): 30-36. [25] Hall EO. The deformation and ageing of mild steel: III discussion of results. Proc Phys Soc London 1951; B64: 747-53. [26] Petch NJ. The cleavage of polycrystals. J Iron Steel Inst 1953; 174: 25-8. [27] Pickering FB. Towards improved toughness and ductility. Kyoto: Climax Molybdenum Company Synopsium; 1971. [28] Calcagnotto M, Ponge D, Raabe D. Effect of grain refinement to 1μm on strength and
toughness of dual-phase steels[J]. Materials Science and Engineering: A, 2010, 527(29): 7832-7840. [29] Wang C, Wang M, Shi J, et al. Effect of microstructure refinement on the strength and toughness of low alloy martensitic steel[J]. JOURNAL OF MATERIALS SCIENCE AND TECHNOLOGY-SHENYANG-, 2007, 23(5): 659. [30] Takeda K, Nakada N, Tsuchiyama T, et al. Effect of interstitial elements on Hall-Petch coefficient of ferritic iron[J]. ISIJ international, 2008, 48(8): 1122-1125. [31] Cao R, Li J, Liu D S, et al. Micromechanism of Decrease of Impact Toughness in Coarse-Grain Heat-Affected Zone of HSLA Steel with Increasing Welding Heat Input[J]. Metallurgical and Materials Transactions A, 2015, 46(7): 2999-3014. [32] Ghosh A, Ray A, Chakrabarti D, et al. Cleavage initiation in steel: Competition between large grains and large particles[J]. Materials Science and Engineering: A, 2013, 561: 126-135. [33] Chakrabarti D, Strangwood M, Davis C. Effect of bimodal grain size distribution on scatter in toughness[J]. Metallurgical and Materials Transactions A, 2009, 40(4): 780-795.
PII: DOI: Reference:
S0921-5093(16)30717-1 http://dx.doi.org/10.1016/j.msea.2016.06.059 MSA33805
To appear in: Materials Science & Engineering A Received date: 26 May 2016 Revised date: 19 June 2016 Accepted date: 20 June 2016 Cite this article as: Xin-li Wen, Zhen Mei, Bo Jiang, Li-chong Zhang and Yazheng Liu, Effect of normalizing temperature on microstructure and mechanical properties of a Nb-V microalloyed large forging steel, Materials Science & Engineering A, http://dx.doi.org/10.1016/j.msea.2016.06.059 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Effect of normalizing temperature on microstructure and mechanical properties of a Nb-V microalloyed large forging steel Xin-li Wen, Zhen Mei, Bo Jiang, Li-chong Zhang, Ya-zheng Liu* School of Material Science and Engineering, University of Science and Technology Beijing, Beijing 100083, China * Corresponding author. [email protected] Abstract The microstructure of a microalloyed large forging steel with different normalizing temperatures ranging from 820 ºC to 940 ºC were characterized. The evolution of austenite formation was determined in a range of heating temperature from 730 ºC to 940 ºC. The mechanical properties were evaluated by tensile test and Charpy V-notch impact test. The relationship between the microstructure and the properties was discussed. The results indicated that the microstructure composed of fine-grained layers (FGL) and coarse-grained layers (CGL) was obtained at 820 ºC. The finest and most homogeneous microstructure and optimal comprehensive mechanical properties were obtained at the normalizing temperature 880 ºC. There was a Hall-Petch relationship between the yield strength and the average grain size, and a linear relationship between the impact energy(KV2) and the reciprocal of the square root of the grain size (D-1/2). Both the strength and toughness of the steel can be attributed to grain refinement. Keywords: Normalizing temperature, Microstructure, Mechanical properties, Nb-V microalloyed, Large forging steel, grain refinement.
1. Introduction Microalloyed large forging steel has been widely used for engineering components which require high strength, good toughness and large size. Large forging steel is usually made from large casting ingots. The as-forged microstructure of the steel is generally composed of coarse ferrite and banded pearlite which consequently lead to limited impact toughness. Due to the characteristic of forging process and the large volume of forging, conventional technology such as thermo-mechanical treatment or accelerated cooling is infeasible for grain refinement in order to improve toughness. However, it is likely to increase toughness levels of microalloyed large forging steel through grain refinement by normalizing. Zhao et al [1] studied the effect of W addition and normalizing conditions on microstructure and mechanical properties of microalloyed forging steels, four kinds of microalloyed forging steels were produced by varying W additions(0, 0.5, 1 and 2 wt%), heat treatment was carried out at temperatures ranging from 840 ºC to 950 ºC followed by air and furnace cooling, the results showed that the microstructure and mechanical properties of the microalloyed forging steels were closely related to the W content, normalizing temperature and cooling method after normalizing. Zhao et al [2] studied the effect of hot forging, normalizing temperature (840 ºC-950 ºC) and cooling method (air and furnace cooling) after normalizing on the toughness and tensile properties of a microalloyed cast steel, the results showed that remarkable improvement in toughness and tensile properties can be obtained by hot forging, proper normalizing temperature and air cooling after normalizing. Zhao et al [3] studied the effect of normalizing temperature (950 ºC-1200 ºC) and cooling method (furnace, air and water cooling) after normalizing on the toughness and tensile properties of a low-carbon microalloyed cast steel, the results showed that heat treatment at 1100 ºC for 2 hours followed by furnace cooling leaded
to the best combination of excellent Charpy impact and tensile properties. The above research focused on the low carbon Nb-Ti microalloyed steel, there was no many studies on Nb-V microalloyed steel in which the V content can be as much as 0.1 wt%. Besides, the sizes of samples used by Zhao et al [1-3] were 11 ×11 × 60 mm3 or 11 ×11 × 110 mm3, the air cooling rate was above 3ºC/s. There were rare studies on large section forging steel with the diameter larger than 250 mm and the air cooling rate below 0.05ºC/s. Austenite formation in low carbon steels has been studied extensively in the literature starting from different microstructures [4-6]. Previous work has shown that in ferritic-pearlitic microstructures the formation of austenite was described as taking place in three main successive steps: (1) nucleation of austenite in pearlite colonies, ferrite-pearlite grain boundaries or ferrite-ferrite grain boundaries, (2) rapid growth of austenite consuming pearlite, (3) slower growth of austenite consuming ferrite [7,8]. Based on the above theory, as for ferrite-pearlite banded microstructure in large forging steel, austenization in ferrite bands and pearlite bands are asynchronous. Nonetheless, rare literature has studied this phenomenon. What’s more, there were no many studies on the effect of intercritical normalizing on microstructure and mechanical properties of Nb-V microalloyed large forging steel. For large forging steel, it is impracticable to enhance cooling rate in case of thermal stress-cracking. The cooling method followed normalizing is usually air cooling, hence the austenization temperature is the decisive normalizing parameter for microstructure and properties. The study of this paper aims at investigating the effect of normalizing temperature on microstructure and properties of a Nb-V microalloyed large forging steel. The evolutions of austenite, microstructure and precipitations of the tested steel were characterized. The relations between the microstructure and properties were discussed. The tested steel in this work with the diameter of Φ290 mm and the V content up to 0.095 wt% has never been investigated. 2. Experimental material and procedures The steel used in this work is a commercial HSLA steel. The chemical composition is listed in Table 1. The round bar specimens with a length of 200mm and diameter 290 mm for normalizing is shown in Fig. 1. They were cut from a Φ290 mm round forging. In order to study the characteristic of microstructure at the corresponding normalizing temperature, cubic samples for quenching were wire-cut from the 1/2 radius of the Φ290 mm round forging. The size of the quenching samples is 10 ×10 × 12 mm3. Both the normalizing and quenching process was conducted in a 45 kw box resistor-stove, the schedules of the process are given in Fig. 2. The normalizing specimens were reheated at 820 ºC, 850 ºC, 880 ºC, 910 ºC and 940 ºC with soaking for 2 hours, respectively, and then were cooled by air with about a 0.03 ºC/s cooling rate. In order to study the evolution of austenite, samples for interrupted heating by quenching were respectively reheated at 730 ºC, 760 ºC, 790 ºC, 820 ºC, 850 ºC, 880 ºC, 910 ºC and 940 ºC holding for 2 hours. After the normalizing process, blanks for metallographic observation and mechanical property test were wire-cut from the 1/2 radius of the normalizing samples along longitudinal axial direction as shown Fig. 1. Metallographic observation direction for all test samples was parallel to the longitudinal section of the Φ290 mm round forging. The microstructure of the samples was etched by a 4% nital solution. The size and area fraction of ferrite and pearlite constituent were measured by software Image-Pro Plus. For each specimen, at least 5 fields of view containing at least 400 grains were measured. Thin foils for Transmission Electron Microscopy (TEM) were prepared using the twin-jet method and observed in a JEM-2100 transmission electron microscope.
The average size and fraction of precipitate particles were statistically measured by averaging 5 fields of view containing at least 300 particles from the images of TEM. In order to study the orientation characteristic of acicular pearlite, electron back scattered diffraction examinations were performed on a field emission gun scanning electron microscope. The blanks for tensile test were machined into standard tensile test specimens of 10 mm in gage diameter and 50 mm in gage length. Tensile tests based on standard of ISO 6892-1: 2009 were carried out on a WDW-200D tensile testing machine at room temperature with a cross-head speed of 0.25 mm min-1 [9]. The yield strength was determined by the 0.2% offset flow stress. All results were repeated for three times and the average values were taken to describe the tensile properties of the test steel. The Charpy V-notched specimens with cross section of 10 × 10 mm2, length of 55 mm, notch angle of 45° and notch depth of 2 mm were employed to study the -40 ºC impact fracture toughness on a ZBC2452-B impact testing machine according to ISO 148-1: 2006. Table 1 Chemical compositions of the tested steel (wt%). C
Si
Mn
P
S
Nb
V
Ti
0.15
0.27
1.45
0.01
0.004
0.045
0.095
0.01
200 mm
Mechanical sample
Metallographic sample 290 mm
Fig. 1 The normalizing sample and sampling method T=820 ºC, 850 ºC, 880 ºC, 910 ºC, 940 ºC
Temperature (ºC)
t=2h
Air cooling
10 ºC /min
about 0.03 ºC /s
Quenching Time (s)
Fig. 2 Normalizing and quenching process 3. Results and discussion 3.1. The evolution of austenite formation The study of the austenite formation was carried out using a DIL805A high resolution dilatometer. Cylindrical samples of 4 mm diameter and 10 mm length were used for the experiments. As shown in Fig. 3, the relations between heating temperature(T) and expansion amount(ΔL) were analyzed to determine Ac1 and Ac3 at a constant heating rate of 10 ºC/min. Since
some investigations have experimentally shown that a separation can be made between the pearlite to austenite and the ferrite to austenite transformation [10-12], an attempt was made to determine the temperature (Acθ) at which this occurred. The determination of Acθ can be only attempted if the first contraction is perceived. It is less evident to see from the dilatometric curve, but easier to determine from the first derivative (dΔL/dT) as shown in Fig. 3. In previous papers studied similar steels, the authors showed that this first contraction was related to the dissolution of pearlite[13, 14]. As shown in Fig. 4(b) and 6(b), microstructure morphologies support the fact that the first contraction observed in the dilatometric curve is due to pearlite dissolution.
Fig. 3 Dilatometric curve and the first derivative According to the results of dilatometer tests, the Ac1 and Ac3 of the tested steel were determined as 725 ºC and 861 ºC, respectively. Fig. 4 shows microstructures obtained after interrupted heating by quenching at 730-940 ºC. It can be seen that the specimen heated at 730 ºC mainly consists of white band and black band microstructure. With increasing temperature from 730 ºC to 850 ºC, the black band microstructure gets wider and wider, the white band gets narrower and narrower accordingly. There is no obvious banded microstructure in samples heated at 880, 910, and 940 ºC. (a)
(b)
(c)
(e)
(f)
Initial pearlite
Initial ferrite (d)
(g)
(h)
Fig. 4 Optical micrographs of quenched microstructure after heating at different temperatures (a) 730 ºC, (b) 760 ºC, (c) 790 ºC, (d) 820 ºC, (e) 850 ºC, (f) 880 ºC, (g) 910 ºC, (h) 940 ºC Fig. 5 shows the SEM micrographs of the quenched microstructure. Fig. 5(a)-Fig. 5(e) corresponds to the micrographs of the white band microstructure in Fig. 4(a)-Fig. 4(e). Fig. 5(f)-Fig. 5(h) are typical micrographs corresponding to Fig. 4(f)-Fig. 4(h). It can be seen from Fig. 5(a) that the microstructure mainly consists of ferrite as well as a small amount of martensite which are distributed at the ferrite boundaries. Formation of martensite in ferrite band indicates that actual austenization has started at 730 ºC. With increasing temperature to 760 ºC, net-like martensite can be seen clearly along the ferrite boundaries, which shows that ferrite boundaries is the preferential site for austenite nucleation and growth. With further increasing of temperature from 760 ºC to 790 ºC, the net-like martensite grows into the ferrite. Accordingly, the amount of ferrite is reduced remarkably. Fig. 5(d) and Fig. 5(e) show that the microstructure is composed of martensite and a small amount of ferrite after quenching at 820 ºC and 850 ºC, respectively. The acicular microstructure adjacent to the ferrite is turned out to be martensite as shown in Fig. 5(i). The microstructure completely consists of martensite when the heating temperature is 880 ºC, 910 ºC, and 940 ºC, indicating that the actual austenization of ferrite band has been finished at about 880 ºC. The formation of acicular microstructures (finger-type austenite) during austenization was also observed by many researchers [15-18]. It is quite possible that the formation of the fingers coincide with the position of former cementite plates that were perpendicular to the grain boundaries. (a)
(b)
ferrite
martensite
(c)
ferrite martensite (d)
martensite
ferrite
martensite
(e)
(f)
ferrite ferrite
(g)
martensite (h)
martensite (i)
Fig. 5 SEM micrographs of quenched microstructure after heating at different temperatures (a) 730 ºC, (b) 760 ºC, (c) 790 ºC, (d) 820 ºC, (e) 850 ºC, (f) 880 ºC, (g) 910 ºC, (h) 940 ºC, (i) micrographs of the acicular microstructure(TEM) Fig. 6 shows the SEM microstructures of the quenched microstructure. Fig. 6(a)-Fig. 6(e) correspond to the micrographs of the black band microstructure in Fig. 4(a)-Fig. 4(e). Fig. 6(f)-Fig. 7(h) are typical micrographs corresponding to Fig. 4(f)-Fig. 4(h). It can be seen from Fig. 6(a) that the microstructure consists of pearlite and ferrite, as well as a small amount of martensite which is distributed at the ferrite-pearlite boundaries. With increasing temperature to 760 ºC, blocky martensite can be seen clearly around ferrite, and pearlite has also been dissolved simultaneously. Fig. 6(c) and Fig. 6(d) show that the microstructure is composed of martensite and a small amount of ferrite after quenching at 790 ºC or 820 ºC. The microstructure completely consists of martensite when the heating temperature is 850, 880, 910, or 940 ºC, indicating that the actual austenization of the pearlite bands has been finished at 850 ºC. (a)
(b)
(c) ferrite ferrite
martensite pearlite ferrite
martensite
(d)
(e)
(f)
ferrite
(g)
(h)
Fig. 6 SEM micrographs of quenched microstructures after heating at different temperatures (a) 730 ºC, (b) 760 ºC, (c) 790 ºC, (d) 820 ºC, (e) 850 ºC, (f) 880 ºC, (g) 910 ºC, (h) 940 ºC Under the given conditions in this paper, the formation of austenite in initial ferrite and pearlite was asynchronous. The nucleation and growth of austenite in initial pearlite was earlier and faster than that in ferrite. The microstructure heredity of austenite on ferrite and pearlite will be shown below. 3.2. Microstructure of normalized samples Fig. 7 shows the optical micrographs of the samples before and after normalizing. As can be seen in Fig. 7(a), as-forged microstructure consists of alternate bands of proeutectoid ferrite and pearlite. EPMA was used to investigate the chemical characteristic across the banded microstructure in Fig. 7(a). Figs. 8(a) and 8(b) show manganese and silicon profiles, respectively. Excellent correlation between microstructural banding and microchemical banding is apparent. In other words, solute lean regions and solute rich regions are consistently associated with regions of proeutectoid ferrite and pearlite, respectively. The results are in full agreement with the findings of
Thompson and Howell [19]. They showed that microchemical banding of manganese was the cause of the microstructural banding observed in a similar steel. (a)
(b)
FGL
(c)
CGL
(d)
(e)
(f)
Fig. 7 The optical micrographs for the specimens before and after normalizing (a) as-forged, (b) 820 ºC, (c) 850 ºC, (d) 880 ºC, (e) 910 ºC, (f) 940 ºC
Fig. 8 Concentration profiles from as-forged sample (a) manganese profile, (b) silicon profile In an attempt to determine the reliability of the data presented in Fig. 8(a), the maximum and minimum concentrations of manganese were estimated using the Scheil equation [20], which was developed for the limiting case of no diffusion in the solid, but complete mixing in the liquid. Cs = kC0(1- fs)k-1 (1) where Cs is the concentration of solute in the solid at a given fraction solidified fs, C0 is the concentration of solute in the alloy, and k is the equilibrium partition ratio, defined as k = Cs/C1 (2) where C1 is the concentration of solute in the liquid in equilibrium with the solid of concentration Cs. To simplify the analysis, k is assumed to be constant throughout the solidification range. For manganese segregation k=0.71 and C0 is 1.45% Mn [21]. Thus, Cs=1.06%Mn for fs=0.1, and Cs=2.00%Mn for fs =0.9. These values are in reasonably good agreement with the maximum and minimum values of manganese concentration shown in Fig. 8(a). Fig. 7(b) and Fig. 9(a) show the inhomogeneous banded microstructure which consists of fine grain bands (FGB) and coarse grain bands (CGB). The FGB is composed of relatively large size of ferrite (initial ferrite) and acicular black microstructure. The CGB consists of relatively fine ferrite (F) and pearlite. The acicular microstructure as shown in Fig. 9(b) is identified as lamellar pearlite by using SEM. Hence, it is called acicular pearlite (AP) in this paper. Color coded inverse
pole figure (IPF) maps of ferrite in CGB is given in Fig. 9(c), which indicates that the adjacent acicular pearlite possesses the same crystal orientation. (a)
FGB
(c)
(b)
AP AP F CGB
Fig. 9 The microstructures obtained with the normalizing temperature 820 ºC (a) OM , (b)SEM, (c) IPF map From the above, complete austenization was finished at about 880 ºC, so there is coarse initial ferrite in the samples normalized at 820 ºC and 850 ºC. When the temperature is 880 ºC, the microstructure is the finest. With increasing temperature from 880 ºC to 940 ºC, abnormal growth of some pearlite nodules occurs, as shown in Fig. 7(e) and (f). The average size of the pearlite nodules in the specimens normalized at 910 ºC and 940 ºC are about 38 and 44 μm in diameter, respectively. These values are very similar to the largest austenite grain sizes in the specimens. As shown in Fig. 10(a), the largest austenite grain size is about 30 μm in diameter in specimen reaustenitized for 2 hours at 880 ºC. The number density is very low. Conversely, as shown in Fig. 10(b) and (c), numerous large austenite grains with the diameter ranging from 40 to 60 μm were observed in the specimens reaustenitized for 2 hours at 910 ºC and 940 ºC. (a)
(b)
(c)
Fig. 10 The micrographs for the specimens reaustenitized at different temperatures (a) 880 ºC, (b) 910 ºC, (c) 940 ºC The similarity in sizes and number densities for the large pearlite nodules and large austenite grains strongly suggests that the large pearlite nodules form within large austenite grains. If it is assumed that air cooling (about 0.03 ºC/s) produces microstructure that are relatively close to equilibrium, and it is further assumed that the pearlite nodules form from a single austenite grain, then the size of the nodules can be estimated from the austenite grain size and the volume fraction of pearlite. If, for simplicity, the austenite grains and the pearlite nodules are assumed to be spherical, then for a pearlite volume fraction of 0.39 and an austenite grain diameter of 60 μm, the diameter of the pearlite nodules would be about 44 μm. Fig. 7(f) and Fig. 10(c) show that large pearlite nodules are about 30-50 μm in diameter, consistent with formation within the largest austenite grains. Hence, it can be concluded that large, irregular pearlite nodules form in abnormally large austenite grains [22]. 3.3. Precipitation Fig. 11 shows the bright field TEM micrographs and EDS analysis of the representative precipitations observed in the as-forged specimen. The observed precipitations can be classified
into three different size ranges: 1-3 μm (cuboidal precipitations), 50-150 nm (spherical precipitations) and 5-15 nm (fine precipitations). The cuboidal precipitations are titanium-rich and niobium-contained (Ti, Nb)C (Fig. 12a and d). The spherical precipitations are niobium-rich and titanium-contained (Nb, Ti)C (Fig. 12b and e) and present a chain-like distribution. The fine precipitations are vanadium-rich and niobium-contained (V, Nb)C (Fig. 12c and f). The majority of the precipitations are duplex-type carbides. Statistics shows that there is no obvious difference of the precipitations in size and distribution between the specimens before and after normalizing. Fig. 12 shows the bright field TEM micrographs of the representative precipitations observed in the normalized specimens.
(b)
(a)
(c)
(Nb, Ti)C (V, Nb)C (Ti, Nb)C
(d)
(e)
(f)
Fig. 11 Three typical TEM micrographs and EDS analysis of the precipitations (a and d)cuboidal precipitation, (b and e)spherical precipitations, (c and f)fine precipitations (a)
(b)
(c)
(Nb, V)C
(V, Nb)C
(Ti, Nb)C
Fig. 12 Three typical TEM micrographs of the precipitations in the specimens normalized at 850 ºC (a)cuboidal precipitation, (b)spherical precipitations, (c)fine precipitations 3.3. Mechanical properties Performance requirements for the tested steel are listed here: Rp0.2≥265 MPa, Rm=450-600 MPa, A≥17%, KV2≥27 J(-40 ºC). Mechanical properties of the tested steels are ineligible before normalizing as listed: Rp0.2=326 MPa, Rm=520 MPa, A=25%, KV2=15J. The heat treated specimens exhibited superior mechanical properties over the original untreated specimen in terms of higher tensile strength, elongation and impact energy as summarized in Table 2.
Table 2 Mechanical properties of the specimens varied with normalizing temperature. T(ºC)
Rp0.2(MPa)
Rm(MPa)
A(%)
-40 ºC KV2 (J)
820
340
540
27
34
850
360
569
32
95
880
372
577
33
123
910
352
567
30
60
940
348
558
29
53
4. Discussion 4.1. Effect of microstructure refinement on yield strength It can be seen in Table 2 that mechanical properties are dramatically influenced by the normalizing temperature. With the increase of normalizing temperature, yield strength (Rp0.2), tensile strength (Rm) and elongation (A) first rise and then drop. The optimum tensile properties are obtained at the normalizing temperature 880 ºC. It is well known that the yield strength for low carbon microalloyed steels is influenced by different strengthening mechanisms, which include solid solution strengthening, dislocation strengthening, fine-grain strengthening and precipitation strengthening [23]. These factors contribute to the yield strength of steel in varying degrees, depending on the microstructure generated by the specific processing technology. Compared with as-forged samples, the yield strength of normalized samples increases 20-83 MPa, This phenomenon is considered to be explained by combination effects of several strengthening mechanisms. Precipitation strengthening can be calculated by Ashby-Orowan equation [24]:
p
0.538Gbf 1/ 2 x ln x 2b
(2)
Where σp is the increase in yield strength (MPa), G is the shear modulus (81600 MPa for Fe), b is the Burgers vector (0.248 nm), f is the volume fraction of particles, and x is the real (spatial) diameter of the precipitates (nm). Using the value of the area fraction and size distribution of precipitates, the precipitation strengthening of samples can be evaluated as Table 3. The subscript 1 represents the fine precipitations and the subscript 2 represents the spherical precipitations. The precipitation strengthening of (Ti, Nb)C can be ignored based on the fact that the large size is not very useful in increasing the strength. Compared with the as-forged samples, the calculated yield strength increments of the normalized samples due to precipitation strengthening are estimated about -7 MPa (820 ºC), -8 MPa (850 ºC), 1 MPa (880 ºC), -2 MPa (910 ºC) and 5 MPa (940 ºC), respectively, as shown in Fig. 13. The actual yield strength increments according to Table 2 are 14 MPa (820 ºC), 34 MPa (850 ºC), 46 MPa (880 ºC), 26 MPa (910 ºC) and 22 MPa (940 ºC), respectively. Comparison of the calculated and actual yield strength increments is shown in Fig. 13, it can be seen that there is no strong correlations between the yield strength change and precipitation strengthening. So the yield strength change cannot be attributed to precipitation strengthening.
Table 3 Average precipitate size, precipitation volume fraction and corresponding precipitation strengthening value of the test specimens. T(ºC)
x1(nm)
f1
σp1(MPa)
x2(nm)
f2
σp2(MPa)
σp1+σp2(MPa)
820
10.6
0.00092
95
112.2
0.00032
9
105
850
11.2
0.00095
93
110.5
0.00037
10
104
880
9.5
0.0009
102
100.6
0.00039
11
113
910
10.5
0.001
100
97.8
0.0003
10
110
940
9.2
0.00094
106
105.4
0.0004
11
117
Fig. 13 Yield strength increment by precipitation Yield strength(Rp0.2) and average grain size(D) of the specimens normalized at different temperatures are shown in Fig. 14. With the increase of normalizing temperature, yield strength first rises and then drops, but average grain size first drops and then rises, indicating that there is a negative correlation between them. The quantitative relationship between the yield strength and grain size has been known as the classic Hall–Petch equation[25,26]： σy=KyD-1/2+σ0 (2) where σy is the yield strength (MPa), σ0 is the frictional stress (MPa), Ky is the Hall–Petch slope, D is the average grain size. According to Eq. (2), fine-grained microstructure has a pronounced effect on the increment of yield strength. In another supporting work, similar conclusion was obtained by Pickering[27] with an empirical equation for ferritic-pearlitic steels under 0.25% C. The method of linear regression is introduced to analyze the relationship between Rp0.2 and D-1/2, the results are shown in Fig. 15. There is a linear relationship between Rp0.2 and D-1/2, so the yield strength change can be mainly attributed to fine-grain strengthening. Similar phenomena can be seen in several papers[28-30].
Fig. 14 Yield strength and grain size of the specimens normalized at different temperatures
Fig. 15 Relationship between yield strength and grain size 4.2. Effect of microstructure refinement on the impact toughness As shown in Table 2, with the increase of normalizing temperature, -40ºC impact energy (KV2) first rises and then drops, the optimum toughness can be obtained at the normalizing temperature 880 ºC. Toughness in steels is controlled by different microstructural constituents that can be considered as material flaws. Some of them such as inclusions are intrinsic while others depend on processing conditions, i.e., boundaries, grain boundary carbides or hard second phases. It can be seen from Fig. 16 that a linear relationship exists between the -40ºC impact energy(KV2) and the reciprocal of the square root of the grain size (D-1/2), indicating the dependence of the impact energy on the average grain size. The impact energy increases with decreasing average grain size. The impact energy is increased nearly by 3 times from 34 J to 123 J when the grain size was refined from 24 μm to 13 μm, indicating that grain refinement contributes largely to the toughness of the steel. The effect of grain refinement on the toughness of the steel can be explained by the generalized Griffith’s equation given below[31-33].
f
2 E (1 v 2 ) D
(3)
where, σf is the cleavage fracture stress, E is the Young’s Modulus, v is the poisson’s ratio, γ is the corresponding effective surface energies for crack propagation across interfaces, D is the grain size.
Fig. 16 Relationship between impact energy and grain size The fractured surfaces of the steel mainly consist of cleavage facets. The representative fracture morphologies of the tested steel normalized at 820 ºC, 880 ºC and 940 ºC are shown in Fig. 17. The arithmetic mean of the observed facet size was determined by linear intercept method on SEM images, the results are given in Table 4. For the specimens with the grain sizes ranging from 24 μm to 13 μm, the arithmetic mean of the cleavage facet size is decreased from
30 μm to 11μm. It implies that the dominant microstructural feature in the toughness of the steel is the grain size. (a)
(b)
(c)
Fig. 17 Fracture surfaces of the impact specimens normalized at (a)820 ºC; (b)880 ºC; (c)940 ºC Table 4 Average grain size and cleavage facet size of the tested specimens. T(ºC)
Grain size(μm)
Cleavage facet size(μm)
820
24
30
850
16
15
880
13
11
910
18
21
940
20
25
Further evidence of the direct relationship between the grain size and the facet size is shown in Fig. 18, which is an IPF map given by EBSD. The figure shows the crack propagation path in the vicinity of the fracture of the broken impact specimen. The different color in Fig. 18 represents the different grain orientations. The black heavy line represents the high-angle grain boundaries(≥15°), the black light line represents the low-angle grain boundaries(＜15°). Fig. 18 shows that the direction of the cleavage crack is largely changed at the high-angle grain boundaries. A local cleavage cracking event across a grain could lead to an instant load drop. However, when the crack attempts to run from one grain to another, the crystallographic orientation and microstructure may change so that more work had to be done to cross the grain boundary. As a result, the local cleavage crack might be arrested [29]. From the above results, it can be concluded that the unit crack path for cleavage fracture is identified as the grain. The grain refinement is very effective in improving the resistance to the cleavage fracture.
Crack
Fig. 18 The crack propagation path of the fractured impact specimen 5. Conclusion The effect of normalizing temperature on microstructure and mechanical properties of a Nb-V-microalloyed large forging steel has been studied in this work. The conclusions are as
follows: (1) When the normalizing temperature was in γ+α dual phase region, the original pearlite band of the tested steel can achieve complete austenization and transformed into relatively fine grain band during air cooling followed, the original ferrite band retained partly and transformed into coarse grain band subsequently. When the normalizing temperature was in γ region, relatively homogeneous ferrite and pearlite grains can be obtained. With the increase of normalizing temperature, both the ferrite and pearlite grain size increase gradually. (2) With the increase of normalizing temperature, strength and toughness of the steel first rise and then drop, the excellent combination of mechanical properties were obtained at the normalizing temperature 880 ºC. The improvement of yield strength and impact energy can be mainly attributed to grain refinement. (3) There is a Hall-Petch relationship between the yield strength and the grain size, and a linear relationship between the impact energy (KV2) and the reciprocal of the square root of the grain size (D-1/2). The strength and impact toughness of the steel can be improved with the refinement in grain size. The unit crack path for cleavage fracture has been identified as the grain. Grain refinement is very effective in improving the resistance to the cleavage fracture. Acknowledgements The authors appreciate the financial support by the National High Technology Research and Development Program (“863”Program) of China (No.2012AA03A503). References [1] Zhao J, Jiang Z, Lee C S. Effects of tungsten addition and heat treatment conditions on microstructure and mechanical properties of microalloyed forging steels[J]. Materials Science and Engineering: A, 2013, 562: 144-151. [2] Zhao J, Jiang Z, Lee C S. Enhancing impact fracture toughness and tensile properties of a microalloyed cast steel by hot forging and post-forging heat treatment processes[J]. Materials & Design, 2013, 47: 227-233. [3] Zhao J, Lee J H, Kim Y W, et al. Enhancing mechanical properties of a low-carbon microalloyed cast steel by controlled heat treatment[J]. Materials Science and Engineering: A, 2013, 559: 427-435.) [4] Li H, Gai K, He L, et al. Non-isothermal phase-transformation kinetics model for evaluating the austenization of 55CrMo steel based on Johnson–Mehl–Avrami equation[J]. Materials & Design, 2016, 92: 731-741. [5] Chang W, Cao W, Yun H A N, et al. Influences of Austenization Temperature and Annealing Time on Duplex Ultrafine Microstructure and Mechanical Properties of Medium Mn Steel[J]. Journal of Iron and Steel Research, International, 2015, 22(1): 42-47. [6] Li S, Kang Y, Zhu G, et al. Austenite formation during intercritical annealing in C-Mn cold-rolled dual phase steel[J]. Journal of Central South University, 2015, 22: 1203-1211. [7] Savran V I, Offerman S E, Sietsma J. Austenite nucleation and growth observed on the level of individual grains by three-dimensional X-ray diffraction microscopy[J]. Metallurgical and Materials Transactions A, 2010, 41(3): 583-591. [8] Souza M M, Guimaraes J R C, Chawla K K. Intercritical austenitization of two Fe-Mn-C steels[J]. Metallurgical Transactions A, 1982, 13(4): 575-579. [9] ISO 6892-1, Metallic Materials-TensileTesting-Part 1: Method of Testat Room Temperature, 2009, ISO E N. 6892-1, Metallic materials–Tensile testing Part 1: Method of test at ambient
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