Materials Performance and Characterization - Forging · PDF file · 2017-07-05in Hot Rolled Steel Plates Reference Homsher, ... m¼the stress of an extrapolated power-function curve

Materials Performance andCharacterization

C. N. Homsher1 and C. J. Van Tyne1

DOI: 10.1520/MPC20150002

Comparison of TwoPhysical Simulation Teststo Determine theNo-RecrystallizationTemperature in HotRolled Steel Plates

VOL. 4 NO. 3 / 2015

Copyright by ASTM Int'l (all rights reserved); Sat Sep 19 17:53:39 EDT 2015Downloaded/printed byCOLORADO SCHOOL OF MINES (COLORADO SCHOOL OF MINES) pursuant to License Agreement. No further reproductions authorized.

C. N. Homsher1 and C. J. Van Tyne1

Comparison of Two PhysicalSimulation Tests to Determine theNo-Recrystallization Temperaturein Hot Rolled Steel Plates

Reference

Homsher, C. N. and Van Tyne, C. J., “Comparison of Two Physical Simulation Tests to

Determine the No-Recrystallization Temperature in Hot Rolled Steel Plates,” Materials

Performance and Characterization, Vol. 4, No. 3, 2015, pp. 1–14, doi:10.1520/

MPC20150002. ISSN 2165-3992

ABSTRACT

Two rolling simulations were conducted using a Gleeble 3500 to determine the

no-recrystallization temperature, TNR on six microalloyed plate steels. Double

hit deformation tests and multistep torsion tests were performed on steels

containing varying amounts of Nb, V, and Ti. TNR for the double hit

deformation tests were determined by finding fractional softening using the

5 % true-strain method and the intersection of the sigmoidal fractional

softening curve with 20 % fractional softening. TNR for the multistep hot

torsion test were determined using a mean flow stress method and finding the

intersection of the two linear regions. TNR values following multistep hot

torsion testing were lower than values measured after double hit deformation

testing. The decrease in measured TNR values for the torsion tests occurs from

the inherent multiple deformations, resulting in refined grains and an increase

in nucleation sites for recrystallization during the subsequent deformation

steps; thus recrystallization can continue to occur at lower temperatures.

Keywords

steel rolling, no recrystallization temperature, double hit compression tests, multistep hot

torsion tests, microalloy

Manuscript received January 16,

2015; accepted for publication

June 17, 2015; published online

July 24, 2015.

1

Department of Metallurgical and

Materials Engineering, Colorado

School of Mines, Golden,

CO 80401.

Copyright VC 2015 by ASTM International, 100 Barr Harbor Drive, P.O. Box C700, West Conshohocken, PA 19428-2959 1

Materials Performance and Characterization

doi:10.1520/MPC20150002 Vol. 4 No. 3 / 0000 / available online at www.astm.org


Introduction

High strength low alloy (HSLA) steels commonly use microalloying additions of V,

Nb, and Ti, generally under 0.10wt. %, for use in large diameter pipeline steels,

structural and automotive applications, and transmission towers. Microalloying is a

useful way to increase strength while minimizing plate thickness, and thus weight.

Microalloying helps control grain size by influencing the no-recrystallization

temperature (TNR) or phase transformations during processing and/or through pre-

cipitation strengthening during cooling [1].

TNR can be studied and quantified by a variety of methods including (1) direct

observations such as optical microscopy and electron backscatter diffraction, and (2)

external physical simulation methods such as multistep hot torsion testing, double-

hit deformation testing, and stress relaxation testing, which are based on material

softening calculations.

Direct measurement of the recrystallized fraction can be difficult in microal-

loyed steels. The material may transform during cooling and special etching techni-

ques, often following a low-temperature heat treatment, are necessary to reveal the

prior austenite grains (PAGs) [2–7]. The procedure is tedious and may be impossible

to use on quenched austenite in alloys with low hardenability. Once a procedure for

revealing the PAGs is determined, it is then often difficult to distinguish between the

recrystallized and deformed grains, leading to a level of subjectivity in the analysis

methodology.

Hence, the preferred methods for determining TNR include two physical simula-

tion techniques. Multi-deformation tests under continuous cooling, such as multi-

step hot torsion testing, can help identify TNR directly. This type of test is limited,

as it does not allow a fundamental study of the static recrystallization behavior

between two rolling passes for a given temperature, as the temperature is ever

decreasing. Isothermal deformations tests are also popular techniques for analyzing

recrystallization behavior. The most common isothermal tests are double-hit and

stress relaxation. These two tests determine the recrystallized fraction as a function

of temperature in the time interval between or after deformation steps. However,

many methods are reported in the literature to evaluate the softening ratio [8–14],

which influences the separation point between softening due to recovery and soften-

ing due to recrystallization. The results from these tests are therefore open to

ambiguity.

Extensive data on recrystallization kinetics are available in the literature;

however, limited information is available comparing data from various testing tech-

niques and analysis methodologies. Gomez et al. [15] and Vervynckt [16] correlated

isothermal double-hit deformation tests with continuous cooling multistep hot tor-

sion testing for static recrystallization kinetics and precipitation interaction, respec-

tively. Maccagno et al. [17] investigated TNR from industrial rolling mills with

laboratory simulations. Although various studies have been conducted over the

years, no standard method exists for determining TNR causing the values to vary

from one investigation to another. However, each method provides insight into the

influence of certain elements and processing parameters on TNR, i.e., precipitation

interaction, strain-rate dependence, etc. Therefore, comparison of data obtained

from different methods and studies must be assessed with precautions to account for

HOMSHER AND VAN TYNE ON PHYSICAL SIMULATION TESTS 2



discrepancies. The current study focuses on comparing double hit deformation test-

ing with multistep hot torsion testing, all performed on a Gleeble 35002.

Experimental Procedures

EXPERIMENTAL MATERIAL

Material for the current study was laboratory-produced, hot-rolled microalloyed

plate steel. The laboratory heats were Nb-microalloyed plate steel to meet API X-70

specifications if processed correctly. Table 1 gives the chemical composition of the six

alloys. The carbon content was held constant at roughly 0.065wt. % and a base Nb

level of 0.060 wt. %. The alloys have varying levels of V, Nb, and Ti with a low and

high level. The Hi-Nb alloy is the control alloy.

DOUBLE HIT DEFORMATION TESTING

Double-hit deformation tests use cylindrical specimens in an axisymmetric compres-

sion test. The test involves reheating to ensure that most precipitates dissolve back

into solution, cooling to deformation temperature (Tdef), compressing to a given

strain (eÞ and with a given strain rate ( _e), holding for an interpass time (tip), deform-

ing the specimen again while holding everything else constant, and determining the

percentage recrystallized or fraction of softening (FS) [12,13,18–22]. The current

study used samples with 10mm diameter and 15mm length, a true strain of 0.2, and

a strain rate of 5 s�1 for all tests. Deformation temperatures (Tdef) were

750�C–1200�C at 50�C increments. The deformation parameters were as follows:

• Soak at austenitizing temperature of 1250�C for 10min• Cool to first deformation temperature of 1200�C at a constant cooling rate of

1.25�C/s• Deform with-e¼ 0.2, _e¼ 5 s�1

• Hold for tip¼ 5 s• Deform with-e¼ 0.2, _e¼ 5 s�1

• Repeat test with new sample and next Tdef

Figure 1 shows a temperature-time schematic of the double-hit deformation test.

The difference between the deformation curves is calculated by finding the fraction

softening (FS). The general equation for FS is given by [9,10,14,19,21–25]:

TABLE 1

Chemical compositions of laboratory Nb-bearing microalloyed steels in wt %.

Material Identification C Mn Si Ti Nb V Al N S P

Lo-Nb 0.063 1.47 0.019 0.006 0.027 <0.001 0.030 0.0041 0.0017 0.012

Hi-Nb 0.066 1.46 0.020 0.007 0.060 <0.001 0.028 0.0039 0.0017 0.011

Lo-V 0.065 1.46 0.016 0.005 0.060 0.021 0.030 0.0046 0.0017 0.012

Hi-V 0.068 1.46 0.017 0.005 0.061 0.056 0.029 0.0040 0.0017 0.012

Lo-Ti 0.062 1.48 0.018 0.028 0.060 <0.001 0.032 0.0050 0.0018 0.011

Hi-Ti 0.065 1.48 0.019 0.099 0.059 <0.001 0.030 0.0040 0.0019 0.011

2Dynamic Systems Inc. (DSI), Poestenkill, NY.




FS ¼ rm � rr

rm � r0(1)

where:

r0¼ the stress at 5 % true strain of the first deformation step,

rr¼ the stress at 5 % true strain of the second deformation step, and

rm¼ the stress of an extrapolated power-function curve at 5 % true strain of the

second deformation step.

The power function simulates the extrapolated curve without any softening.

Figure 2 shows a double-hit deformation stress-strain plot, extracting specific points

to calculate FS. In the current study, each test was run in triplicates to ensure

repeatability.

MULTISTEP HOT TORSION TESTING

Multistep hot torsion tests simulate the rolling process through a series of deforma-

tion steps (i.e., “hits”) and continuous cooling for a given set of parameters, such as

e, _e, tip, and temperature range. The torque and the amount of twist are measured

and converted into stress and strain. One sample is used for the entire temperature

range and is deformed through the TNR. The current study used samples with

10mm diameter and 20mm gauge length, a true strain of 0.2, and a strain rate of

5 s�1 for all tests. Deformation temperatures were 750�C–1200�C at 25�C incre-

ments. The deformation parameters were as follows:

• Soak at austenitizing temperature of 1250�C for 10min• Cool to first deformation temperature of 1200�C at a constant cooling rate of

1.25�C/s• Deform with-e¼ 0.2, _e¼ 5 s�1

• Cool to next Tdef in 20 s at a cooling rate of 1.25�C/s

FIG. 1

General temperature-time

schematic of a double-hit

deformation test.




• Deform with-e¼ 0.2, _e¼ 5 s�1

• Continue test until the final Tdef (750�C) is reached

Figure 3 shows temperature-time schematic for multistep hot torsion tests.

Figure 4 shows an example of the data collected from a multiple step hot torsion test.

FIG. 3

General schematic of a

multistep hot torsion test.

FIG. 2

Example of a double-hit

deformation curve used to

determine fractional softening

via the 5 % true-strain method.




From the procedure developed by Richardson et al. [26], the torque-twist data can

be converted into equivalent stress-strain. The Von Mises equivalent stress, r, is

given by,

r ¼ 3ffiffiffi3p

T2pa3

(2)

where:

T¼ torque (N-m), and

a¼ the radius (m) of the gauge section.

Equivalent strain, e, is given by,

e ¼ 0:724ahffiffiffi3p

l(3)

where h is the angle of twist (radians), and

l¼ the length of the gauge section (m).

A value of 0.724 is used as the effective radius following Richardson et al. [26]

and Barraclough et al. [27]. The concept of effective radius to calculate the stress

strain curve mitigates problems associates with the strain gradient across the diame-

ter of the specimen. Figure 5 shows the results of converting the torque-twist data

into stress-strain data. The mean flow stress (MFS) is given by

MFS ¼ 1eb � ea

ðeb

ea

r de(4)

FIG. 4

Example of equivalent torque-

twist data from multiple step

hot torsion test.




where:

eb and ea¼ the final and initial strains per pass, and

r¼ the summation of the stress per pass.

Once the MFS is calculated, it is plotted against the inverse of absolute tempera-

ture. Multiple stages are seen in the MFS graph, and TNR can be determined from

the transition between two of these stages. Figure 6 shows an example of MFS versus

the inverse absolute temperature. TNR is the intersection of the linear fit between

stages I and II. In this example, TNR is 958�C, which is read as 8.123� 10�4 K�1 off

the plot. The linear regions were found using a least squares regression.

Results and Discussion

DOUBLE HIT DEFORMATION TESTING

Figure 7 shows a plot of double-hit deformation true-stress true-strain curves at

1200, 1000, and 750�C for the Lo-V alloy. As expected, the flow stress of the test per-

formed at 1200�C is lower than the tests performed at 1000 and 750�C, while the

flows stress of the test performed at 750�C is higher than the tests at 1000 and

1200�C. Also noticeable in Fig. 7 is the comparison of the two flow curves at each

temperature. The 1200�C curves are almost identical, indicating 100 % fraction soft-

ening (FS). The 1000�C flow curves are similar, with the second curve showing more

work hardening than the first, indicating something less than 100 % FS. The 750�C

curves are similar to an interrupted test, where the second curve appears to be a con-

tinuation of the first, indicating close to 0 % FS. The fraction softening was found

for each alloy at each condition. Figure 8 shows the fraction softening as a function

of temperature for the Lo-V alloy. The circles indicate the average of the three tests

FIG. 5

Example of equivalent stress-

strain calculated from torque-

twist data shown in Fig. 4 from

a multiple step hot-torsion test.




FIG. 7

Plot of double-hit deformation

true-stress true-strain curves at

1200, 1000, and 750�C.

FIG. 6

Example of mean flow stress

versus absolute temperature

calculated from equivalent

stress-strain data.




and the line represents a sigmoidal fit to the data. The dashed line indicates the 20 %

FS, which defines TNR [12].

MULTISTEP HOT TORSION TESTING

Following the analysis procedure provided above, the mean flow stress (MFS)

and subsequently the TNR for the six alloys was determined. Figure 6, above,

shows an example of MFS as a function of inverse absolute temperature for the

Lo-V alloy. Figure 9 shows the MFS versus inverse absolute temperature for the

three Lo-V test specimens on a single graph. The plot shows little variation in

MFS for the three specimens, indicating the testing conditions were consistent

for each run. The average TNR for the three specimens is 953�C with a range of

953�C–954�C.

Table 2 gives the TNR values obtained from double hit compression and

from multistep hot torsion for the six alloys in the current study. The TNR values

for hot torsion tests were expected to be lower than the double-hit tests due to

an expected difference in grain size at the TNR. At the measured TNR, the torsion

samples have already undergone a large number of deformation passes with a

large amount of total deformation. Multiple recrystallization cycles have

occurred, leading to grain refinement in the austenite. The grain size in the dou-

ble hit compression sample prior to the first deformation step is expected to be

large due to the high reheating time and temperature. With the refined grains in

the torsion sample, more nucleation sites are available for recrystallization, lead-

ing to a lower TNR.

FIG. 8

Fractional softening of the Lo-V

alloy. TNR is denoted by the

dotted line at 20 pct. FS which

intersects close to 1000�C.




COMPARING THE TWO TESTING SIMULATIONS

The data follow the predicted outcome, having a lower TNR from multistep hot tor-

sion than from double-hit deformation simulations. However, Fig. 10 shows a weak

trend in the data between hot torsion and double-hit tests. There appears to be no

clear indication of how much influence hot torsion tests have on TNR compared with

double-hit deformation tests, i.e., the difference of TNR between hot torsion and

double-hit simulations is not consistent. Other factors may contribute to the

reported values of TNR. For example, double-hit deformation tests generally only

measure static recrystallization (recrystallization which occurs between passes),

while hot torsion tests may experience dynamic recrystallization (recrystallization

which occurs during deformation) due to a buildup of strain at lower temperatures,

near and below TNR [28]. Specimens in multistep hot torsion tests undergo multiple

deformations, leading to refined grains and a higher likelihood of strain-induced

precipitation. In comparison, each specimen in a double-hit deformation test has

TABLE 2

Comparison of TNR determination through double-hit compression and multistep hot torsion.

Material Identification Double-Hit (�C) Torsion (�C) DT (�C)

Lo-Nb 953 931 22

Hi-Nb 1000 960 40

Lo-V 995 953 42

Hi-V 1025 954 71

Lo-Ti 981 959 22

Hi-Ti 1026 994 32

FIG. 9

Comparison of the three Lo-V

test specimens for MFS as a

function of inverse absolute

temperature. The three tests

were very similar, showing

consistency in testing

procedure.




a long hold at high temperatures, generating larger austenite grains. Any

strain-induced precipitation would only occur in the temperature regime for

strain-induced precipitates, and not in every sample. The increased grain boundary

area in the torsion tests provides more nucleation sites for recrystallization at lower

temperatures; thus driving TNR to lower values. Other considerations include the

temperature for the double-hit deformation tests were controlled using type-K ther-

mocouples, while a dual-frequency pyrometer was used for the hot torsion tests.

There is some evidence in the literature to suggest a “break-down” of the type-K

thermocouples at high temperatures for long holds [29]. It is not believed to be the

main contributing factor in the weak trend between hot torsion and double-hit sim-

ulations. Other possibilities may be due to inconsistencies in test setup, such as

ensuring isothermal tests, use of nickel-paste, tantalum foil, and grafoil to provide a

barrier to carbon diffusion from the tungsten carbide platens and test sample,

extended period of time necessary to run the tests, extending over weeks, resulting

in various opportunities for inconsistencies.

The thermocouples used in the double-hit deformation tests have an uncertainty

of 65�C. Additional uncertainty can be attributed to the analysis methodology. The

pyrometer used in the multistep hot-torsion tests also has an uncertainty of 65�C.

Since a more direct method was used to determine TNR for the torsion test, being

able to determine TNR from a single test, not multiple tests, the uncertainty of the

test is less than that of the double-hit deformation study. The average difference of

calculated TNR values of the multistep hot torsion tests for a given alloy is 2�C, while

double-hit deformation tests have a difference of about 15�C. The variation in TNR

range may be attributed to the analysis method of the two testing procedures. The

double-hit tests were analyzed following the 5 % true strain method to determine

fraction softening. This method required an extrapolated power function line of the

FIG. 10

Comparison of the

experimental TNR for multistep

hot torsion tests and double-hit

deformation tests. The strain

and strain rate for each test

was 0.2 and 5 s�1, respectively.




first stress-strain curve as well as the stress value at 5 % true-strain for both curves.

The hot torsion tests were analyzed by determining the mean flow stress. The area

under each stress-strain curve was determined and normalized by dividing the stress

by the amount of strain per pass. The total stress per pass normalized by pass strain

appears to lead to a more robust analysis procedure.

Conclusions

An overview of various physical simulation methods for determining the no-

recrystallization temperature, TNR, was described, focusing on comparing results

from double hit deformation and multistep hot torsion tests to simulate a hot rolling

process. The double hit deformation tests result in higher values for TNR than those

measured from multistep hot torsion testing. The difference is primarily due to the

different grain sizes since multistep hot torsion testing undergoes multiple deforma-

tion passes prior to and after TNR, where double hit deformation tests undergo

deformation at a single temperature after a reheat and cool cycle. The smaller grains

in multistep hot torsion tests lead to more nucleation sites for recrystallization, and

thus a lower TNR.

ACKNOWLEDGMENTS

The writers would like to thank the Colorado School of Mines for supporting this

research through the Advanced Steel Processing and Products Research Center.

References

[1] Gladman, T., The Physical Metallurgy of Microalloyed Steels, 2nd ed., Maney

Publishing, London, 2002.

[2] Barraclough, D. R., “Etching of Prior Austenite Grain Boundaries in

Martensite,”Metallography, Vol. 6, No. 6, 1973, pp. 465–472.

[3] Brewer, A. W., Erven, K. A., and Krauss, G., “Etching and Image Analysis of

Prior Austenite Grain Boundaries in Hardened Steels,” Mater. Charact.,

Vol. 27, No. 1, 1991, pp. 53–56.

[4] Chapman, B. H., Cooke, M. A., and Thompson, S. W., “Austenite Grain Size

Refinement by Thermal Cycling of a Low-Carbon, Copper-Containing Marten-

sitic Steel,” Scr. Metall. Mater., Vol. 26, No. 10, 1992, pp. 1547–1552.

[5] Bodnar, R. L., McGraw, V. E., and Brandemarte, A. V., “Technique for Reveal-

ing Prior Austenite Grain Boundaries in CrMoV Turbine Rotor Steel,” Metal-

lography, Vol. 17, No. 1, 1984, pp. 109–114.

[6] Garcia de Andres, C., Caballers, F. G., Capdevila, C., and San Martin, D.,

“Revealing Austenite Grain Boundaries by Thermal Etching: Advantages and

Disadvantages,”Mater. Charact., Vol. 49, No. 2, 2003, pp. 121–127.

[7] Zhang, L. and Guo, D. C., “A General Etchant for Revealing Prior-Austenite

Grain Boundaries in Steels,”Mater. Charact., Vol. 30, No. 4, 1993, pp. 299–305.

[8] Vervynckt, S., Verbeken, K., Lopez, B., and Jonas, J. J., “Modern HSLA Steels

and Role of Non-Recrystallisation Temperature,” Int. Mater. Rev., Vol. 57,

No. 4, 2012, pp. 187–207.




http://dx.doi.org/10.1016/0026-0800(73)90044-X

http://dx.doi.org/10.1016/1044-5803(91)90079-J

http://dx.doi.org/10.1016/0956-716X(92)90254-C

http://dx.doi.org/10.1016/0026-0800(84)90009-0

http://dx.doi.org/10.1016/0026-0800(84)90009-0

http://dx.doi.org/10.1016/S1044-5803(03)00002-0

http://dx.doi.org/10.1016/1044-5803(93)90078-A

http://dx.doi.org/10.1179/1743280411Y.0000000013

[9] McQueen, H. J. and Jonas, J. J., “Role of the Dynamic and Static Softening

Mechanisms in Multistage Hot Working,” J. Appl. Metalwork., Vol. 3, No. 4,

1985, pp. 410–420.

[10] McQueen, H. J., “Review of Simulations of Multistage Hot-Forming of Steels,”

Can. Metall. Q., Vol. 21, No. 4, 1982, pp. 445–460.

[11] Luton, M. J., Dorvel, R., and Petkovic, R. A., “Interaction Between Deformation,

Recrystallization and Precipitation in Niobium Steels,” Metall. Trans., Vol. 11,

No. 3, 1980, pp. 411–420.

[12] Palmiere, E. J., Garcia, C. I., and DeArdo, A. J., “The Influence of Niobium

Supersaturation in Austenite on the Static Recrystallization Behavior of Low

Carbon Microalloyed Steels,” Metall. Mater. Trans. A, Vol. 27, No. 4, 1996,

pp. 951–960.

[13] Kwon, O. and DeArdo, A. J., “Interactions Between Recrystallization and Pre-

cipitation in Hot-Deformed Microalloyed Steels,” Acta Metall. Mater., Vol. 39,

No. 4, 1991, pp. 529–538.

[14] Rao, K. P., Prasad, Y. K. D. V., and Hawbolt, E. B., “Study of Fractional Soften-

ing in Multi-Stage Hot Deformation,” J. Mater. Process. Technol., Vol. 77,

Nos. 1–3, 1998, pp. 166–174.

[15] Gomez, M., Rancel, L., Fernandez, B. J., and Medina, S. F., “Evolution of

Austenite Static Recrystallization and Grain Size During Hot Rolling of a

V-Microalloyed Steel,” Mater. Sci. Eng. A, Vol. 501, Nos. 1–2, 2009, pp.

188–196.

[16] Vervynckt, S., Verbeken, K., Thibaux, P., Liebeherr, M., and Houbaert, Y.,

“Austenite Recrystallization–Precipitation Interaction in Niobium Microalloyed

Steels,” ISIJ Int., Vol. 49, No. 6, 2009, pp. 911–920.

[17] Maccagno, T. M., Jonas, J. J., Yue, S., McCrady, B. J., Slobodian, R., and Deeks,

D., “Determination of Recyrstallization Stop Temperature From Rolling Mill

Logs and Comparison With Laboratory Simulation Results,” ISIJ Int., Vol. 34,

No. 11, 1994, pp. 917–922.

[18] Vervynckt, S., Verbeken, K., Thibaux, P., Liebeherr, M., and Houbaert, Y.,

“Control of the Austenite Recrystallization in Niobium Microalloyed Steels,”

Mater. Sci. Forum, Vols. 638–642, 2010, pp. 3567–3572.

[19] DeArdo, A. J., “Niobium in Modern Steels,” Int. Mater. Rev., Vol. 48, No. 6,

2003, pp. 371–402.

[20] Wilber, G. A., Bell, J. R., Bucher, J. H., and Childs, W. J., “The Determination

of Rapid Recrystallization Rates of Austenite and the Temperatures of

Hot Deformation,” Trans. Metall. Soc. AIME, Vol. 242, No. 11, 1968,

pp. 2305–2308.

[21] Devadas, C., Samarasekera, I. V., and Hawbolt, E. B., “The Thermal and Metal-

lurgical State of Steel Strip During Hot Rolling: Part III. Microstructural

Evolution,” Metall. Trans., Vol. 22, No. 2, 1991, pp. 335–349.

[22] Vervynckt, S., 2010, “Control of the Non-Recrystallization Temperature in

High Strength Low Alloy (HSLA) Steels,” Ph.D. thesis, Universiteit Gent, Gent,

Belgium.

[23] Le Bon, A. B. and de Saint-Martin, L. N., “Using Laboratory Simulations to

Improve Rolling Schedules and Equipment,” Micro Alloy., Vol. 75, 1975,

pp. 90–98.




http://dx.doi.org/10.1007/BF02833663

http://dx.doi.org/10.1179/cmq.1982.21.4.445

http://dx.doi.org/10.1007/BF02654565

http://dx.doi.org/10.1007/BF02649763

http://dx.doi.org/10.1016/0956-7151(91)90121-G

http://dx.doi.org/10.1016/S0924-0136(97)00414-7

http://dx.doi.org/10.1016/j.msea.2008.09.074

http://dx.doi.org/10.2355/isijinternational.49.911

http://dx.doi.org/10.2355/isijinternational.34.917

www.scientific.net/MSF.638-642.3567

http://dx.doi.org/10.1179/095066003225008833

http://dx.doi.org/10.1007/BF02656802

[24] Laasraoui, A. and Jonas, J. J., “Prediction of Steel Flow Stresses at High Temper-

atures and Strain Rates,” Metall. Trans. A, Vol. 22, No. 7, 1991, pp. 1545–1558.

[25] Poliak, E., “Recrystallization During Hot Rolling,” presented at the Austenite

Processing Symposium (Internal Company Presentation), September 18, 2008

-unpublished.

[26] Richardson, G. J., Hawkins, D. N., and Sellars, C. M., Worked Examples in

Metalworking, The Institute of Metals, London, 1985.

[27] Barraclough, D. RH. J., Nair, K. D., and Sellars, C. M., “Effect of Specimen

Geometry on Hot Torsion Test Results for Solid and Tubular Specimens,”

J. Test. Eval., Vol. 1, No. 3, 1973, pp. 220–226.

[28] Solhjoo, S. and Ebrahimi, R., “Prediction of No-Recrystallization Temperature

by Simulation of Multi-Pass Flow Stress Curves From Single-Pass Curves,”

J. Mater. Sci., Vol. 45, No. 21, 2010, pp. 5960–5966.

[29] Dynamic Systems Inc., Gleeble Systems Application Note: Diffusion Effects on

Type K (Cr-Al) Thermocouple Measurements, DSI: Dynamic Systems Inc.,

Poestenkill, NY, 2001.




http://dx.doi.org/10.1007/BF02667368

http://dx.doi.org/10.1520/JTE10007J

http://dx.doi.org/10.1007/s10853-010-4681-3

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Materials Performance and Characterization - Forging · PDF file · 2017-07-05in Hot Rolled Steel Plates Reference Homsher, ... m¼the stress of an extrapolated power-function curve