perlitik ray çeliği

8/2/2019 perlitik ray elii

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Satyam S. Sahay,1

Goutam Mohapatra,2

and George E. Totten3

Overview of Pearlitic Rail Steel: Accelerated Cooling,Quenching, Microstructure, and Mechanical Properties

ABSTRACT: Railway networks form an integral part of the infrastructure development of a developingcountry with ever-increasing passenger and freight volume. Increase in train speed, pay load, reliability, andsafety are the major thrust areas for railways requiring more stringent mechanical properties such as wear,deformation resistance, and fatigue life from the railway steel. Steel chemistry control and thermomechani-cal processing significantly affect final properties and performance of the railway steel. For example, for agiven steel composition, a number of stable or metastable microstructures can be obtained by controllingheat treatment operations. Conventional rail steels primarily contain nearly eutectoid pearlitic microstruc-ture, which is dependent on the criticality of the application. An overview of the physical metallurgy prin-ciples involved during the manufacturing of rail steel will be provided here. The primary focus of this reviewis thermal processing including quenching and accelerated cooling of the rail steel. In addition, otherimportant aspects relating to design and production of rail steel are discussed, including: impact of steel

chemistry on the phase diagram, effect of thermomechanical processing on microstructure, and influenceof microstructure or residual stress on mechanical properties.

KEYWORDS: accelerated cooling, rail steel, pearlitic steel, quenching, cooling curves,microstructure, residual stress, mechanical properties

Introduction

For over 150 years, railway infrastructure has continued to be one of the important backbones for passen-

ger as well as freight transportation around the globe. Increases in train speed, pay load, reliability, and

safety are the major thrust areas for railways, which necessitates stringent mechanical and functional

properties such as wear and deformation resistance, fatigue life from the railway steel. These ever-

increasing performance targets have been successfully met by the manufacturers of rail steel. Interestingly,

rail steel provides a classical case for teaching physical metallurgy principles. This includes a clear

understanding of steel chemistry dependence on transformation temperatures, influence of thermomechani-

cal processing on microstructure, and direct bearing of microstructure features such as pearlite lamellae

spacing on mechanical properties as well as functional performance parameters. More often than not, for

successful commercial production of rail steel, metallurgists have worked backward in the above process

chain, i.e., based on functional and mechanical property requirements and microstructure specifications.

Practical constraints of cooling rates and available time for conducting the required thermomechanical

processing have necessitated the design of a transformation diagram by tuning the alloy chemistry.

A review of the physical metallurgy principles involved during the manufacturing of rail steel is

provided here. The primary focus is on thermal processing including quenching and accelerated cooling of

the rail steel. In addition, other important aspects for the design and production of rail steel are discussedincluding: impact of steel chemistry on the phase diagram, effect of thermomechanical processing on

microstructure, and influence of microstructure or residual stress on mechanical properties. Although all

three common grades of rail steelpearlitic, ferritic, and bainitic steelsare used for railway applications

including track and wheels, this review will primarily be limited to pearlitic steel for railway track. Recent

Manuscript received July 15, 2008; accepted for publication May 11, 2009; published online June 2009.1

Tata Research Development and Design Centre, Tata Consultancy Services Ltd., 54, Hadapsar Industrial Estate, Pune 411 013,

India, e-mail: [email protected]

Tata Research Development and Design Centre, Tata Consultancy Services Ltd., 54, Hadapsar Industrial Estate, Pune 411 013,

India.3

Dept. of Mechanical and Materials Engineering, Portland State Univ., Portland, OR 97201, e-mail: [email protected]

Journal of ASTM International, Vol. 6, No. 7

Paper ID JAI102021

Available online at www.astm.org

Copyright 2009 by ASTM International, 100 Barr Harbor Drive, PO Box C700, West Conshohocken, PA 19428-2959.


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developments of bainitic rail steels with excellent mechanical properties and wear resistance 1,2 will be

discussed in a subsequent review.

Chemistry and Phase Diagrams

Continuous cooling transformation diagrams are very important in designing the required thermomechani-

cal processes of steel. This can be illustrated using a transformation diagram Fig. 1 for the eutectoid steel

3. In this figure, the pearlite, bainite, and martensite start Ps, Bs, Ms and end Pf, Bf, Mf temperatures

are marked. By superimposing the cooling pathways on this transformation diagram, the phases expectedfrom the transformation can be identified. For example, at very slow cooling rates path A, full pearlitic

microstructure is expected. As the cooling rate is increased path C, the cooling path crosses pearlitic and

bainitic transformation boundaries yielding a mixed microstructure. If a very fast cooling path is under-

taken path D, it is possible to completely avoid the pearlitic and bainitic transformation boundaries,

leading to martensitic structure. It must be noted that from the rail steel perspective, none of these three

paths are highly desirable. Cooling path A leads to coarse pearlitic structure with low hardness and wear

resistance, path C leads to highly undesirable mixed pearlitic and bainitic microstructure having low wear

resistance, and path D leads to brittle martensitic structure. The highly desired fine pearlitic microstructure

associated with excellent hardness, wear resistance, and fatigue properties can be obtained by following

cooling path B where the component is rapidly cooled to a lower pearlitic start temperature that is located

just above the bainitic start temperature and a nearly constant thermal transformation is conducted up to

the pearlite finish boundary, which is followed by the cooling to ambient temperature. Cooling path B is

commonly referred to as interrupted cooling or accelerated cooling and is used extensively for rail steel

processing.

In practice, the constant temperature pearlitic transformation can be achieved simply by cooling

rapidly by airwater mist spray to the pearlitic start temperature and then applying high cooling rates in

small zones followed by slow air-cooling regimes. The slow air-cooling regimes prevent material from

crossing the bainitic boundary. The cooling process and the heat of transformation recalescence effect

produce a nearly uniform constant temperature profile throughout the cross section of the rail.

As explained above, for a given material, the desired microstructure can be obtained by controlling the

cooling rates. However, it is possible that the required cooling rate for a given component is too high to be

achieved in practice. Alternately, if the component cross section is very large, the core of the component

may cool at the desired rate. In either of these situations, it is possible to shift the phase boundary by alloyaddition. This has been illustrated by comparing the transformation diagram of two rail steels with minor

FIG. 1Effect of cooling profile on the final microstructure (adapted from Ref [2]).

2 JOURNAL OF ASTM INTERNATIONAL


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differences in their chemistry. The two steels considered in this example contain equal C 0.76, Si 0.25,

P 0.017, and S 0.014 content but minor differences in other trace elements as shown in Table 1 3. Except

for Mn content, the difference in the other trace elements is very low. In spite of such minor differences,

the continuous cooling transformation diagram given in Fig. 2 shows significant differences for the two

alloys. The pearlite and bainite phase boundary for alloy B with higher trace element content has consid-

erably shifted toward higher time relative to alloy A. A cooling rate of 350C /min will result in coarse

pearlite phase for the alloy A, whereas it barely touches the bainite start boundary, resulting in primarily

martensitic microstructure 3. This example illustrates the importance of tight chemistry control during

rail steel processing. Fortunately, during the last few decades, significant advances in secondary steelmak-

ing and casting have resulted in very clean steels.

Some of the commonly used pearlitic and bainitic steels have been tabulated in Table 2 2. Both

silicon and manganese are well-known solid solution strengtheners of ferrite 4. Manganese reduces both

the interlamellar spacing of pearlite and the prior austenite grain size, both of which can be beneficial to

strength and toughness. Silicon addition of 0.20 % increased the yield strength YS about 10 %.

TABLE 1The trace elements in the two rail steels.

Mn Ni Cr Mo V Al Sn Sb As Cu

Alloy A 0.81 0.015 0.02 0.02 0.02 0.003 0.002 0.0008 0.006 0.033

Alloy B 1.14 0.09 0.21 0.021 0.02 0.007 0.02 0.005 0.009 0.26

TABLE 2Compositions of some of the typical pearlitic and bainitic rail steels [2].

Number C Si Mn Ni Mo Cr V Nb B Al Ti Type

1 0.55 0.25 1.0 Pearlitic

2 0.8 0.3 1.0 Pearlitic

3 0.7 1.9 1.5 Pearlitic

4 0.75 0.7 1.0 1.0 0.1 Pearlitic

5 0.65 0.25 0.7 Pearlitic

6 0.04 0.2 0.75 2.0 0.25 2.8 0.01 0.03 0.03 Bainitic

7 0.09 0.2 1.0 0.5 0.003 0.03 0.03 Bainitic

8 0.07 0.3 4.5 0.5 0.1 Bainitic

9 0.1 0.3 0.6 4.0 0.6 1.7 0.01 0.03 0.03 Bainitic

10 0.3 0.2 2.0 0.5 1.0 0.003 0.03 0.03 Bainitic

11 0.3 1.0 0.7 0.2 2.7 0.1 Bainitic12 0.52 0.25 0.35 1.5 0.25 1.7 0.1 0.01 Bainitic

FIG. 2Effect of steel grades on phase transformation during continuous cooling (adapted from Ref [2]).

SAHAY ET AL. ON OVERVIEW OF PEARLITIC RAIL STEEL 3


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Manganese addition of0.35 % increased strength by about 67 % 4. The addition of a small amount of

niobium to rail steel considerably improves the wear resistance, weldability, ductility, and fracture tough-

ness 4,5. Ductility and fracture toughness are important when the steel is to be used in cold environments

with heavy-haul iron ore service. The role of niobium in head-hardened rail steel is to refine the prior

austenite grain size and prevent grain growth ensuring the formation of fine pearlite colonies 4,5.

An addition of about 0.08 % by weight vanadium increases the YS to approximately 20 % over

nonvanadium alloys of the same composition. Vanadium contributed to the strength in these alloys prima-

rily by solid solution strengthening and also, to some extent, by grain refinement. The addition of vana-

dium also produces the undesirable effect of increasing the Charpy transition temperature. This is due to

an undesirable rolling texture verified by x-ray diffraction. Chromium addition to grade 1100 steel shiftsthe continuous cooling transformation curve to the right and, as a result, even still-air cooling after

hot-rolling results in 100 % pearlitic structure 4,5.

Hydrogen has a deleterious influence such as flaking or shatter cracks on rail steel if present above 3

ppm 6. One way to control hydrogen content in the liquid steel is by adopting advanced secondary steel

making e.g., arc argon oxygen decarburization to reduce the hydrogen content to an acceptable limit

below 3 ppm. The high wear resistance THS 11 grade rail steel developed by Thyssen Mill, Germany

contains around 3 ppm of oxygen by adopting a vacuum degassing technique 7.

Another way to control the hydrogen content is by slow or retarded cooling, since slow cooling gives

enough time for hydrogen to diffuse out of the steel. It was also observed that the response of critical

amounts of hydrogen to form flakes depends upon alloy content, for example, increase in sensitivity to

hydrogen increases with an increasing manganese-carbon ratio 7. Higher sulfur contents reduce suscep-

tibility to the formation of flakes. A chromium-manganese special grade is less sensitive to hydrogenflaking. The nonmetallic inclusion content also depends upon oxygen content in the steel. Vacuum degas-

sing is used to reduce the oxygen content to 0.0020.008 %. The percentage of sulfur prints with inclu-

sions reduced from 26 to 1 between 1978 and 1986 8.

Quenching

Heat Transfer

Quenching plays an important role in the thermomechanical processing of steel, which can be illustrated

by the following example. Consider a typical CMn rail steel containing 0.77 % C, 0.95 % Mn, 0.22 % Si,

and 0.1 % Cr. The required cooling rates under continuous cooling condition for this steel to obtain

various types of microstructures are shown in Table 3. In contrast, to obtain the desired fine pearlitic

microstructure, an interrupted cooling profile is required where the steel is cooled from 1050 to 580C at

the rate of 640C /min, followed by slow cooling at 45C /min to 480C, before giving moderate cooling

at the rate of 380C /min to room temperature 3.

There are three primary heat transfer cooling mechanisms, most commonly observed during conven-

tional immersion quenching in a vaporizable quenchant such as water. When water first contacts the hot

metal surface, it is surrounded by a vapor film where heat transfer occurs by full-film boiling FB. When

the temperature decreases to the Leidenfrost temperature, the vapor film collapses and surface wetting by

the liquid occurs by a nucleate boiling NB process 9. When the surface temperature decreases further

to a temperature less than the boiling point, NB ceases and convective cooling CONV begins.

Each of these cooling mechanisms is associated with different heat transfer processes as illustrated inFig. 3 10. This is important because of the significantly different heat transfer coefficients HTCs

TABLE 3Influence of cooling rate on the microstructure of rail steel [3].

Cooling rate Phases

240C/min Pearlite

250C/min Pearlite+ bainite

400C/min Pearlite+ bainite+ martensite

643C/min Martensite

Interrupted cooling Fine pearlite



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corresponding to these cooling processes. For example, when water is used to quench steel, typical HTCs

are FB FB 100250 W /m2 K, NB NB 1020 kW/m

2 K, and CONV CONV ca. 700 W /m2 K.

The simultaneous presence and relative stability of these widely varying heat transfer conditions are an

important factor influencing nonuniform cooling and increased stresses during a water quenching process11,12.

The total amount of heat transferred to the quenchant heat transfer is denoted by the symbol Q.When heat is released from the hot body to the cooler surroundings, Q is negative. The heat flow per unit

time heat transfer rate is measured in watts and is quantitatively defined as

Q =dQ

dt1

Heat flux is defined as the heat transfer rate per unit of cross-sectional area q is measured as watts per

square meter and is determined experimentally by measuring the temperature difference over a material

with a known thermal conductivity. However, measured cooling rates are dependent on the size, shape, and

thermal properties of the steel. A common method of quantifying the cooling properties of a quenchant thatis to calculate the HTC is

FIG. 3Illustration of cooling curve and the different cooling mechanisms occurring from the time of

initial immersion until the conclusion of the quench.



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dQ

dt = Tp Tc A 2

where:

=HTC,

TP =surface temperature of the steel,

TC=temperature of the quenchant,

A =surface area, and

dQ /dT=heat flux amount of heat transferred.

The HTC relates to quenching and cooling and is classically related to the Grossman value H, whichis equal to the HTC divided by two times the thermal conductivity H= /2. When the H-value

is multiplied by the diameter of the body D, the product corresponds to the dimensionless Biot number

Bi, which relates to the resistance to heat transfer both at the surface and inside of a body

HD =Bi=

R 3

The physical significance of the Biot number can be understood by considering the heat flow from a hot

metal to the quenchant during the cooling process. The factors inhibiting heat flow are: a heat flow from

the metal to the surface and b heat flow from the surface into the quenchant. This equation means that

heat transfer is proportional to section size thickness of the metal being quenched and that the HTC at the

interface between the cooling metal and the quenchant is inversely proportional to the thermal conductivity

of the metal. To more accurately relate the Biot number to size and shape, the generalized Biot criterion,

BiV is calculated from

BiV=

VL =

VK S

V4

where:

L = characteristic length, which is commonly defined as the surface area of the body S divided by the

volume of the body V and

K= Kondratyev form coefficient shape factor. A limited number of these values is provided in Table

4 13.

The generalized Biot criterion BiV is related to the temperature difference within the steel and the

cooling metal surface to the quenchant by the dimensionless Kondratyev number Kn. There is universal

correlation between Kn and generalized Biot number BiV and it is useful for any configuration of steel part.

The Kondratyev number may be calculated as follows or it may be obtained from Table 5 14:

Kn = BiV =BiV

BiV2+ 1.437BiV+ 1

1/25

is the temperature field nonuniformity criterion, which is equal to

=Tsf Tm

TV Tm6

where:

Tsf=average temperature of the surface of the component being quenched,

Tm =temperature of the quenchant, and

TV=average temperature over the volume of the component.The value can also be defined in terms of the generalized Biot criterion Bi V

TABLE 4Equations for calculation of Kondratyev shape factor K for simple shapes.

Shape of Body K S /V

Parallelepiped with sides L1, L2, L3 L12 +L2

2 +L32 /2 2L1

1 +L21 +L3

1

Cylinder of infinite size with height Z 5.783R2 +9.87Z21 2R1 +Z1

Sphere R2 /2 3 /R



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= 1BiV

2+ 1.437BiV

2+ 11/2

7

Kobasko et al. 14 and later Fernandes and Prabhu 15 used these relationships to calculate the HTC

from the two temperatures T1 and T2 corresponding to the times t1 and t2 taken from the quenchant

cooling rate curve. The cooling rate m is then calculated from

m =lnT1 Tm lnT2 Tm

t2 t18

The Kondratyev number Kn is calculated from the cooling rate m

Kn =m K

9

For a cylindrical body, the Kondratyev form factor K is R2 /5.783. The HTC for film boiling FB iscalculated from

FB =BiV V

KA10

In addition to simplified methods for the calculation of HTC such as those described here, one may also

use inverse analysis such as that described by Beck 16. This methodology is commonly performed on

cooling curve data 17 and has a number of advantages over those methods described above since surface

heat flux can be calculated from known positions of the heated problem. However, these methods havebeen described by others and will not be discussed further here 18.

TABLE 5Summary of generalized Biot numbers, values, and Kondratyev numbers.

BiV Kn fBiV 1 / 1 / BiV0.00 1.000 0.000 1.000 1.000 1.00

0.01 0.993 0.010 0.005 1.010 1.00

0.10 0.931 0.093 0.040 1.074 0.97

0.20 0.868 0.174 0.070 1.152 0.95

0.30 0.811 0.243 0.092 1.23 0.93

0.40 0.759 0.304 0.11 1.32 0.92

0.50 0.713 0.356 0.124 1.40 0.90

0.60 0.671 0.403 0.135 1.49 0.89

0.70 0.633 0.443 0.146 1.58 0.88

0.80 0.599 0.479 0.154 1.67 0.87

0.90 0.568 0.511 0.160 1.76 0.86

1.00 0.539 0.539 0.170 1.85 0.85

1.50 0.430 0.646 0.190 2.32 0.82

2.00 0.386 0.712 0.210 2.81 0.81

2.50 0.304 0.760 0.213 3.29 0.79

3.00 0.264 0.792 0.220 3.79 0.76

4.00 0.210 0.839 0.230 4.76 0.76

4.50 0.190 0.854 0.231 5.26 0.76

5.00 0.174 0.869 0.233 5.75 0.76

6.00 0.148 0.889 0.236 6.76 0.76

7.00 0.129 0.903 0.240 7.75 0.75

9.00 0.1026 0.929 0.240 9.75 0.75

10.00 0.0931 0.951 0.240 10.74 0.74

20.00 0.0482 0.965 0.240 20.75 0.75

50.00 0.0197 0.986 0.240 50.76 0.76

100.00 0.0099 0.993 0.240 100.30 0.76

0.0000 1.000 0.240



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Quenchants

A variety of fluids, including air, mist, water, oil, and gases have been used as quenchants. The effective-

ness of various quenching media can be compared through their heat transfer rates. As can be seen from

Table 6, still air provides a very low heat transfer rate of 50 80 W /m2 compared to 3000 3500 W /m2for circulated water 19. Heat transfer rates for various gases are shown in Fig. 4. The HTC, a measure of

effectiveness of the cooling media, is determined from the experimentally measured cooling profiles. For

example, the HTC calculated from cooling curves using finite element modeling is shown in Fig. 5 20.

In addition to the quenchant type, the cooling rates achieved by the metal also depend on the operating

parameters, such as bath circulation or temperature. For example, it can be seen from Fig. 6, that minor

changes in bath agitation can significantly modify the cooling profile of the metal through mixing of the

fluid as well as influencing the convective heat transfer 19. Similarly, Fig. 7 shows that significant

changes in cooling rates can be achieved by changing the bath temperature 19. Increasing the bath

temperature reduces the cooling capacity of the quenchant.

Scale on the mill products can also significantly affect the cooling profiles in the metal. The influence

of scale thickness on the central cooling profiles during oil quenching is illustrated in Fig. 8 19. It mustbe noted that scale in the mill product is not desired since it reduces the yield of the material. In actual

practice, scale thickness is controlled by tightly controlling the air-fuel ratio, temperature, and mill delays

in the reheating operation 21. It must be noted in Fig. 8, that a very thick nonadherent scale reduces the

cooling rate due to trapped air gap whereas very thin adherent scale can enhance the cooling rate due to

increased surface roughness. Measurement of surface temperature is also important since it is often used as

one of the boundary conditions in mathematical models to estimate the core temperature 21.

TABLE 6Comparison of typical heat transfer rates [19].

Quench Medium

Heat Transfer

W /m2

Still air 5080

Nitrogen 1 bar 100150

Salt bath or fluidized bed 350500

Nitrogen 10 bar 400500

Helium 10 bar 550600

Helium 20 bar 9001000

Still oil 10001500

Hydrogen 20 bar 12501350

Circulated oil 18002200

Hydrogen 40 bar 21002300

Circulated water 30003500

FIG. 4Comparison of HTCs achieved with different gas quenching media (adapted from Ref [19]).



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Different quenching media are used in the heat treatment of rail steel. Traditionally, air, water, oil, or

mixtures such as air and water mist are also used as quenchants. Although water provides the highest

cooling rate, it is not suitable for all types of steels. Mist quenching or fog quenching is often used in rail

steel production and involves rapid heat extraction from a metal by a fast-moving stream of gas, usually

air, that contains water droplets. The cooling capacity of the fog is due to the absorption of heat by the gas

and the heat of vaporization of the water. The addition of water droplets fog to an air stream can

significantly increase its cooling capacity 22. For comparison, forced air convection exhibits a maximum

heat flux between 2 3 W /cm2 whereas an airwater mist, depending on the mass flux and atmospheric

conditions, is capable of exhibiting a maximum heat flux of 100 1200 W /cm2 23.

Mist fog quenching is performed using with a mixed spray jets gas typically air and a liquidtypically water. The diameter of the water droplet in the mixed spray jet may vary from approximately

10400 m 2427. The mixing ratio of the gas and liquid may be either constant, as in conventional

fog quenching, or it may be varied from pure air to pure water, as required, throughout the cooling process.

In conventional fog quenching, the cooling capacity is regulated by the distance between the spray nozzle

FIG. 5The HTC as a function surface temperature for the head (dotted line) and flange (dashed line)

obtained by employing inverse analysis (adapted from Ref [20]).

FIG. 6Effect of bath agitation on the cooling rate (adapted from Ref [19]).



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and the object to be quenched 28. The spray jet that is used to control the injection of the liquid into the

gas may utilize a hydraulic or a pneumatic atomizer. A pneumatic atomizer is reported to provide smaller

droplet sizes with a more uniform distribution 27.

Heat transfer of the mist is dependent on the Leidenfrost temperature TL. If the surface temperature

of the steel is greater than TL during the mist cooling process, the surface of the steel will be covered by

a vapor blanket of water vapor 28. And if the surface temperature of the steel is less than TL, the vapor

blanket will have collapsed and the surface will be wetted by the water droplets with a corresponding

increase in heat transfer. The Leidenfrost temperature is related to the spray orifice, flow rate, physical and

chemical properties of the liquid, spray pressure, and momentum of the liquid droplets. Droplet momen-

tum may be increased by increasing the liquid droplet size or increasing the impact velocity.

The composition of the spray has a significant effect on the droplet size and droplet speed andtherefore droplet kinetic energy during heat transfer. The ability of a spray droplet to come into contact

with the surface is dependent on the vapor layer that is formed. Direct contact produces higher heat

transfer rates. The kinetic energy of the droplet is quantified by the Weber number, which is a dimension-

less number used to analyze multiphase fluid flows such as spray droplets of water in air. It is a measure

FIG. 7Effect of bath temperature on the cooling rate (adapted from Ref [19]).

FIG. 8Effect of scale thickness on the cooling rate (adapted from Ref [19]).



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of the relative importance inertia of the fluid relative to its surface tension. The Weber number is useful in

modeling droplet contact heat transfer and is the governing parameter for contacting droplet deformation.

The Weber number We is calculated from 27

We=v2l

11

where:

=fluid mist density kg /m3,

v=velocity of the droplet m/s,

d=diameter of the droplet m, and=surface tension of the fluid dyne/cm.

A low Weber number can be increased most easily in practice by increasing the droplet velocity. In

some cases, a critical Weber number is reported, which is defined as the lowest value sufficient to complete

the breakup of liquid droplets in a high speed gas stream. When designing a quenching process, it is often

desirable to maximize the time that the surface temperature is greater than the Leidenfrost temperature to

minimize the cooling rate to allow the steel to become sufficiently rigid to resist plastic deformation, which

may lead to increased distortion.

When water alone is used as a quenching medium for low-carbon steel, it results in strain development

and distortion or crack formation. Sometimes a mixture of water and air mist is used as a quenching

medium 29. A recently developed quenching technology using compressed air single step or com-

pressed air followed by mist two step instead of only mist mixture of water and air developed by Zhan

and Wang 30 ensures a very stable cooling speed, which remains unaffected by surface condition, and

thus ensures stable properties of the rail.

Another alternative to using water as a quenchant is the use of an aqueous polymer solution. Typically,

when the hot steel is initially immersed into the quenching fluid, a vapor film surrounds the hot metal

surface, resulting in film boiling. As the surface temperature decreases, NB begins and continues until the

surface temperature is less than the boiling point of the quenching fluid, at which point CONV begins. The

non-steady-state behavior obtained when hot metal is quenched into water and oil, is shown in Figs. 9 and

10, respectively, where it is observed that all three cooling mechanisms occur on the surface simulta-

neously. This is significant because the HTCs for these cooling processes vary by an order of magni-

tude or greater, thus producing substantial thermal gradients during the cooling process that lead to

increased residual stresses 9.

Because of the different wetting phases on the metal surface and the enormous differences in thevalues ofFB, NB, and CONV, the time-dependent temperature distribution within the metal specimen

FIG. 9Wetting process of a cylindrical CrNi-steel specimen 25-mm diameter100 mm quenchedfrom 850 C into distilled water at 30 C with an agitation rate of 0.3 m/s.



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will also be influenced by the velocity and geometry of the wetting front as well as the geometry of the

quenched part. One way to minimize the magnitude and effect of these thermal gradients is to use

agitation. Ju et al. 31 have shown that agitation facilitates film breakage and decreases the thickness of

the insulating steam film during film boiling.

A quenchant patented by Blackwood and Cheeseman 32 comprises a mixture of water and polyalky-

lene glycol PAG 219 parts; remainder is water. The PAG ensures more uniform heat transfer and thus

reduces the internal stress and subsequent distortion of the work piece. Many aqueous polymer solutions,

such as the 10 % solution of the PAG quenchant used for this work, exhibit a pseudo-Newtonian cooling

process. This means that the cooling mechanism is essentially the same all over the surface of the cooling

part at any point in time. This is illustrated in Fig. 11 19. In this case, the hot metal surface is covered

by a uniform polymer film assuming uniform agitation separated from the hot metal by a vapor filmbarrier. As the specimen is cooled, the polymer film explosively ruptures and rapid cooling results by NB.

When the NB process ends, the polymer is redissolved into the solution and the part is cooled by convec-

FIG. 10Wetting process of a cylindrical CrNi-steel specimen 25-mm diameter100 mm quenched

from 860 C into petroleum oil at 60 C without agitation.

FIG. 11Wetting process of a cylindrical silver specimen 12.55-mm diameter45 mm quenched from

850C into a 10 % aqueous solution of a PAG solution at 25C without agitation (an agitated specimenwould exhibit substantially more uniform film formation).



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tion. It is important to note that quenching into an aqueous polymer quenchant solution results in a

fundamentally different and more uniform cooling process than when quenching into either water or oil.

The effect of non-Newtonian versus Newtonian wetting on thermal gradients and residual stress

formation was described by Tensi et al. 33 and Narazaki et al. 34. The influence of the non-Newtonian

wetting process on the temperature distribution within the quenched specimen is shown in Fig. 12, where

the temperature measured near the surface of a submerged cylindrical specimen at different locations from

the lower end is shown for water and polymer aqueous solution. If there were an explosionlike wetting of

the specimen, such as would occur in a Newtonian wetting process typically exhibited by a PAG polymer

quenchant, the different cooling curves shown in Fig. 12 would be essentially equivalent. If the tempera-tures were measured in the center of the specimen, the differences in the thermal gradients exhibited by

these two different cooling wetting processes would not be evident 33.

Polyacrylate quenching media for rail steel applications have been reported by Kopietz and Munjat

35 and by Tokuue and Kato 36 who reported the use of aqueous solutions containing 0.410 % of salts

sodium, potassium, or ammonium of polyacrylic acid, polymethacrylic acid, and copolymers of meth-

acrylic acid and acrylic acid with an intrinsic viscosity of 0.010.05 l/g 32. These ionic polymers are

thermally stable, corrosion resistant, and easily waste disposable. The water-soluble polymer typically

coats the hot metal forming an insulating film around the workpiece, which reduces heat transfer rates and

the extended film boiling vapor-blanket cooling favors the formation of pearlite at the expense of

typically undesirable martensite during quenching. In addition, soft spotting and excessive distortion

control and cracking problems are reduced relative to water and oil quenching.Nakamura et al. 37 showed that contamination leads to nonuniformity of surface heat transfer.

Examples of contamination effects leading to quench nonuniformity included water contamination of

quench oil, presence of sludge, and nonuniform film deposition when aqueous polymer quenchants are

used.

Spray Quenching

Lenard 38 has extensively characterized the effect of nozzle type, geometry, and strength of spray on the

cooling rates of the rail by embedded thermocouple experiments. The cyclic nature of accelerated cooling

of the rail profile was simulated under laboratory conditions. Strength of the spray was found to have the

most significant effect on the cooling rates. In this experiment, an initial cooling rate of as high as

148C /s from 925C was observed for nozzles with 65 angles, 0.43-m/s velocity located at 63.5 mmfrom the rail. In another paper, Stewart et al. 39 have discussed the effectiveness of various hydraulic

FIG. 12Cooling curves received for different positions within an instrumented cylindrical probe: (a) The

solid line is for water and (b) the dashed line is for a 5 % aqueous solution of a polyalkylene glycolpolymer quenchant.



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nozzles, such as hollow cone conical shaped water jet only in outer shell of cone, full cone solid-conical

shaped water jet, flat spray solid rectangular shaped water jet, and solid stream on heat removal from hot

metal.

The reduced effectiveness of cooling due to contaminated water, e.g., inadvertent mixing of lubricants,

has been discussed. The possibility of limitation of the cooling rate due to conduction of heat from the core

to the surface and the resulting ineffectiveness of an increase in spray power was illustrated. As can be

seen in Fig. 13, when the HTC is increased by increasing the spray power, in the initial stage a minor

increase in the HTC significantly reduces the cooling time 39. However, when the HTC is increased

beyond 12 kW/m2 K, there is a minor reduction in cooling time. At initial stages, when the HTC is low

the cooling of components is limited by removal of heat from the surface whereas at the later stage cooling

gets limited by the conduction of heat from the core to the surface.

The cooling profiles under various cooling rates for the rail steel have been modeled 40 by finite

element method for the cooling rates of 615C /s. The temperature rise due to recalesence heat released

during phase transformation has been simulated and temperature profiles at various locations have been

predicted. A parametric study on the effect of nozzle width, discharge velocity, and cooling water tem-

perature was carried out on the accelerated cooling of steel strip 41. Nozzle width was found to have the

most significant influence on the thermal behavior of the strip due to significant enhancement in the width

of impingement region resulting in an increase in heat extraction rate. It was noted that the coiling

temperature reduces by 200C when the nozzle width is doubled from 1 to 2 cm 41.

Accelerated Cooling

Accelerated cooling refers to controlled cooling methods, using an airwater mist, polymer, or water

cooling, which are faster than the air cooling. The accelerated cooling was initially used in the hot-strip

mill of Brinsworth, UK, to achieve lower coiling temperature with improved mechanical property without

the necessity of the long cooling zone after the rolling 42. Accelerated cooling has played a significant

role in the production of modern rail steel, as it enables phase transformation at lower temperature

resulting in lower pearlite interlamellar spacing with low wear resistance. For example, for the conven-

tional CMn rail steel, the pearlite interlamellar spacing can be reduced to 100 nm at the nose of the

pearlite start temperature as compared to 200400 nm 3 achieved by the air cooling at higher transfor-

mation temperature. There are essentially three variants of accelerated cooling referred as a conventional

heat treatment, b off-line, and c in-line head hardening processes.

In the conventional heat treatment process, the entire rail is reheated to austenizing temperature840860C and subsequently quenched in oil to produce a fine pearlitic structure 3. However, it is

FIG. 13The time to reach 300C at a point 4 mm below the surface of steel plate from a start tempera-

ture of 500C (adapted from Ref [39]).



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important to account for the increase in oil bath temperature due to heat absorption from hot steel as well

as austenite to pearlite heat of transformation recalescence. This is achieved by oil bath circulation. Also,

the bainite phase is avoided by maintaining the oil at 40C. This methodology was widely used as it

resulted in premium rail steel with very uniform and fine pearlitic microstructure. In the method, the head,

web, and base sections of the rail achieves uniform strength and hardness. For the conventional CMn

steel, the typical Brinnel hardness values that can be achieved by this method at the rail head surface and

core are 375 and 365, respectively 3. The residual stress developed during the quenching is removed by

tempering the rail at around 450C. The major disadvantage of this method is the necessity of theadditional processing step of reheating quenching, resulting in higher process cost as well as reduction in

plant productivity.

In the off-line head-hardening method, the head of the rail is selectively heated to the austenitic

temperatures by induction or flame heating followed by accelerated cooling using airwater mist to obtain

a very fine pearlitic structure. Due to the selected heating, high quenching rates are achieved by the bulk

of the steel, in addition to the external cooling. Due to the selective nature of the process, there is a

significant variation in microstructure and mechanical properties in the rails produced by this method. For

example, for the conventional CMn steel, the Brinell hardness at the top surface could be 368 as

compared to 287 in the core 3. One of the major concerns of this method is the significant difference in

mechanical properties at the junction of head and web, which at times results in separation, especially at

the curved tracks. Also, significant residual stresses are generated during this selective hardening process.

Furthermore, the necessity to reheat, even though selectively, the rail in this method is not desired due to

higher process cost and lower productivity.

The in-line head-hardening process is the most sophisticated and efficient method, where the acceler-

ated cooling process forms an integral part of the hot-rolling process. This method utilizes the heat of the

hot-rolling process, thereby eliminating additional reheating steps. The computerized cooling control con-

tinuously changes the water flow rates to achieve constant microstructure throughout the length of the rail.

As the process keeps pace with the hot-rolling milling, four to five times higher productivity levels are

achieved as compared to the off-line method. The high productivity and low operational cost justifies the

high capital cost associated with precise control system. The precise cooling control system results in very

uniform microstructure and properties. The hardness profiles of the in-line head-hardened process has been

compared with conventional and off-line processes in Fig. 14. It can be seen that the hardness profile very

similar to the conventional method and far superior to the off-line method can be achieved by thistechnique. Another advantage of the in-line method is that the sophisticated cooling produces very straight

FIG. 14Comparison between hardness profiles developed during fully heat treated, off-line, and in-line

hardening of rail sections (adapted from Ref [3]).



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rails, thereby removing additional straightening steps essential for off-line method.

In the in-line method developed by Algoma Steel Corporation, Canada 8 the rail is cyclically cooled

from 790 to 565C. The cyclic cooling is achieved by low pressure water spray followed by air-recovery

zones between the water sprays. The typical composition of this rail steel comprised 0.70.8 % C, 0.75

1.05 % Mn, 0.20.5 % Si, 0.20.7 % Cr, 0.020.1 % V, and 0.05 % Mo. As illustrated in Fig. 15, baffles

are used to prevent water spray from reaching the web region, thereby keeping the rail flange hot, while

the head is force cooled. The temperature time plot for the three different locations across the head of the

rail is shown in Fig. 16. The cyclic cooling profile due to air-recovery zones can be observed near the

surface of the rail head. Also, increase in steel temperature due to heat flow from inside in the recovery

zone as well as heat of transformation can be observed. Furthermore, the nearly constant temperature

profile throughout the cross section, after the initial cooling, can be noted in this method. The hardness

profile from surface to core of the rail head obtained by this method is shown in Fig. 17 for two different

transformation temperatures. It can be observed that fairly uniform hardness can be achieved by this

method due to the precisely controlled cooling condition described above. Furthermore, it can be noted

that significant improvement in hardness can be achieved by reducing the transformation temperature. For

example, when the transformation temperature was reduced from 638 to 582C, the surface hardness

increased from 300 to 340 BHN. This is due to a reduction in pearlite lamellae spacing with a reduction

in transformation temperature 43.

Different companies have developed variants of in-line cooling methods. For example, Nippon Steel

can process several rails in an in-line head-hardening steel, whereas VoestAlpine steel uses dipping of rail

head into aqueous synthetic polymers 3. In a recent work, Zhan and Wang 30 have described significantimprovements in head-hardening technology in China. The new induction heating coil enables higher

FIG. 15A schematic presentation of water spray configuration for in-line head hardening (adapted from

Ref [43]).



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uniformity of 50C as compared to 100C obtained in the conventional method. This results in uniform

hardness in the head hardened rail.

In this innovative two step cooling method 30, first compressed air is used for austenite to pearlite

transformation followed by mist cooling for obtaining straight rail profiles. In the first step, the transfor-

mation is accurately controlled whereas in the second step deformation is controlled. It must be noted that

the introduction of second step reduces the high power consumption associated with the air compressor.

The uniform hardness obtained by this method is shown in Fig. 18. Furthermore, significant improvement

in impact toughness 29.141.8 J /cm2 has been achieved over the in-line method. The rails heat treated

by this methodology are extensively used in China as well as in South Koreas railway network 30.

FIG. 16The cooling profile of rail steel at three different locations in the head portion (adapted from Ref

[43]).

FIG. 17The hardness profile of the rail head measured from the top of the surface to the interior as

shown in the inset picture for the as-received sample and heat treated isothermally at various temperatures(adapted from Ref [43]).



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Microstructure and Mechanical Properties

Austenite to pearlitic transformation is a diffusion-controlled process and is governed by nucleation and

growth processes. Carbon diffusion through austenite is the rate controlling step during this transforma-

tion. Most of the mechanical properties depend on the formation of pearlitic structure. A fine and 100 %

pearlitic structure, low interlamellar spacing, produces improved mechanical properties such as wearresistance, i.e., the smaller the interlamellar spacing, the greater the wear resistance and vice versa. These

microstructures can be tailored by understanding the transformation mechanism. The interlamellar spacing

varies inversely with undercooling below the equilibrium eutectoid transformation temperature; the lower

the transformation temperature, the finer the interlamellar spacing. The effect of transformation tempera-

ture on mechanical properties has been illustrated in Fig. 19.

Rail steels may contain variable microstructures such as grade 700, which possesses a mixed micro-

structure of 30 % ferrite and 70 % coarse pearlite. Grade 1100 exhibits up to 100 % fine pearlitic structure

for head-hardened HH 1100 grades. Their selection is dependent on the required properties. Both grades

1100 and HH 1100 exhibit a tensile strength of 11001200 MPa 44. Grade 1100 rail steel contains around

1 % Cr whereas grade HH 1100 contains no Cr Table 7.

Wear resistance along with good weldability, strength, and fracture resistance of rail steels are amongsome of the essential properties required for heavy axle loads and improved train speed. Conventional

rail steels primarily contain near-eutectoid pearlitic microstructure. Pearlite characteristically possesses a

FIG. 18Comparing the hardness profile of the head-hardened rails cross section (adapted from

Ref [30]).

FIG. 19The effect of cooling stop temperature on YS and UTS of rail steel (adapted from Ref [43]).



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mixture of alternating layers lamellae of relatively soft ferrite and a hard brittle iron carbide called

cementite. The pearlitic structure exhibits good wear resistance because of the presence of hard carbides

and toughness due to the presence of a soft ferrite phase.

Wear behavior of pearlitic rails has been studied at the Facility for Accelerated Service Testing

FAST, in Pueblo, Colorado. FAST has a 4.8-mile long rail, which is subjected to actual train service. In

one study, the wear behavior was determined for two steels exhibiting the same 100 % pearlitic structure,

although one steel possessed a greater hardness than the other. The results showed that wear is a three-

stage process E. L. Brown and G. Krauss, unpublished research, Colorado School of Mines, 1982, where

the first stage involves severe plastic deformation and the depth of the deformed zone was found to be

shallower in the case of the harder steel. The second stage involves formation of subsurface cracks, whichspalled off the rail as small slivers or flakes. The first stage of wear can be controlled by steel hardness. In

study, Clayton and Danks 45 observed that the hardness of pearlite increases with decreasing interlamel-

lar spacing as shown in Fig. 20. They also studied the effect of wear rate on pearlite spacing as well as on

hardness. The results Figs. 2022 show that the wear rate decreases with decreasing interlamellar spacing

and increasing hardness 45.

Hyzak and Bernstein 46 studied the effect of prior austenite grain size and pearlite colony size on

mechanical properties as shown in Fig. 16 for fully pearlitic microstructures see also Fig. 23. They

observed that for the same pearlite spacing a coarse austenite provides better hardness and YS than a fine

austenite grain size. It was observed by Houin et al. 47 that the YS and the ultimate yield strength UTS

increase with increasing carbon content with the eutectoid composition 0.8 %C exhibiting a maximum.

The ductility measured by reduction in an area of eutectoid steel 100 % pearlite was found to besignificantly less than hypoeutectoid steels 47. For the same interlamellar spacing, pearlite is discontinu-

ous in hypoeutectoid steel and continuous in eutectoid steels. However, hypoeutectoid steels exhibit lower

TABLE 7Typical chemical composition, tensile strength, and microstructure of different rail steels [44].

Rail Steels %C %Mn %Cr Tensile Strength MPa Microstructure

Crane rail 0.35 0.8 600 50 % pearlite50 % coarse pearlite

Flat Bottom Rails

Grade 700 0.5 1.0 700 30 % ferrite70 % coarse pearlite

Grade 900 0.75 900 100 % coarse pearlite

Grade 1100 0.75 1.1 0.9 1100 100 % fine pearlite

Head-hardened 1100 0.75 1.0 1100 100 % fine pearlite

FIG. 20Hardness of the pearlite phase as a function of interlamellar spacing for various steels (adaptedfrom Ref [45]).



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YS than eutectoid steels. It was also observed that pearlite with a carbon content of 0.6 % with a very fine

lamellar structure produces high YS as well as a reduction in area.

Rolling fatigue is the predominant failure mechanism in the rail steel. The interplay of residual stresses

and microstructure features on fatigue crack initiation and propagation is extremely complex and neces-

sitates a detailed discussion, which will be addressed in the subsequent article. In the following section, the

effect of microstructure on fatigue failure and processing induced residual stress and the resultant distor-

tion will be briefly covered.

Fatigue defects are mainly caused by maximum shear stresses in the upper part of the rail head and by

stress concentrations near nonmetallic inclusions in the matrix. Moreover, an increase in tensile strength isoften recommended for the improvement of fatigue properties of rail steels 48. Fine pearlite structure is

preferred versus coarse pearlite for the prevention of crack initiation. In addition, fatigue crack propagation

in pearlitic rail steels also depends upon the grain boundary ferrite. Cyclic loading performed 49 on

FIG. 21Wear rate of pearlite as a function of interlamellar spacing for rail steels tested at two differentcontact pressures as indicated (adapted from Ref [45]).

FIG. 22Wear rate of pearlite as a function of hardness for rail steels tested at two different contactpressures as indicated for a series of rail steels (adapted from Ref [45]).



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the single-edge notch specimens machined from the rail heads and wheel webs indicate grain boundary

ferrite in a pearlitic microstructure. The results Fig. 24 show that traces of grain boundary ferrite

24 % by volume may reduce crack growth rates by a factor of two at intermediate and low growthrates 5104 mm/cycle. The railhead contains mostly nonmetallic sulfide stringer inclusions

aligned parallel to the rail axis. Fatigue experiments performed with sulfide inclusions in the rail steel from

0.097 to 0.318 percentages do not exhibit an observable effect on the overall crack growth rate or the

fatigue growth mechanism.

Masumoto et al. 50 report that a quenched and tempered microstructure and a fine pearlite micro-

structure exhibit slower fatigue crack propagation rates than a coarse pearlite microstructure.

Residual Stress and Distortion of Rail Steel

As the cross sections of the railway track steel are fairly large, the thermomechanical processing becomes

nonuniform. During the thermomechanical processing, there are spatial variations in cooling rates acrossthe cross section as well as the length of the rail. As a result, different locations undergo transformation at

FIG. 23(a) Hardness and (b) YS as a function of pearlite interlamellar spacing for fully pearlitic

microstructures (adapted from Ref [46]).



22/26

different time, resulting in residual stresses in the rail. The residual stress developed in the hot rolling

process as well as during the cooling process, leads to bending or fatigue failure before and/or during

service 51. The phase transformation austenite/pearlite or austenite to bainite in a material results in

specific volume change that cannot be elastically accommodated and results in plastic deformation. A

homogeneous temperature distribution produces a homogeneous plastic deformation throughout the work

piece, however, if the temperature is inhomogeneous different temperature at different part of the work

piece some part of the deformation energy is stored as residual stress because different parts of the work

piece undergo different rates of transformation. Therefore, the distortion of rail production line after hot

rolling is due to the inhomogeneous temperature profile within the rail producing different transformation

rates 52.Another important source of residual stress is during the roller straightening operation to correct for

buckling that develops during cooling of the rail. Apart from the development of residual stress during the

rail manufacturing process, bucking can also be caused by the rolling of the wheels at the head area of rails

FIG. 24Fatigue crack growth rate as a function of fatigue cycle for pearlitic rail steel (Ref [49]).

FIG. 25The calculated bending (camber) as function of cooling time for the rail steel (adapted fromRef [53]).



23/26

that comes in contact with the wheels. It is important to estimate the magnitude of distortion and residual

stress during the manufacturing of rail steel. Residual stress development and distortion of rail during

cooling to ambient temperature after hot rolling has been simulated 53. The thermomechanical analysis

was carried out using general purpose ANSYS software. The magnitude of the calculated bending cam-

ber during air cooling to ambient temperature is shown in Fig. 25.

The computed residual stress varies from a 27.5 MPa in tension at the middle of the web to 83.5 MPa

in compression at the tip of the base Fig. 26. Moreover, the calculated stress value indicates a change of

stress from compressive to tensile from the surface to the center of the head and a change from tensile to

compressive stress state from the center to the tip of the base. The variation of residual stress along the

symmetry line is shown in Fig. 26b. The state of residual stress is compressive at the head and tensile at

the web as well as the base.

Sommer et al. 54 patented a heat treatment method wherein the distortion can be reduced substan-

tially. The method comprises preheating the entire rail at around 1000F, which is below the transforma-

tion temperature followed by induction heating the rail head to 1900F and flange at 1500F, which is

above transformation temperature along with a balanced thermal deformation along the neutral axis. The

next step is to spray quench the head and flange to 1000F, which is below the transformation temperature

to produce a desired microstructure while achieving a balanced thermal deformation of the rail about the

neutral axes. In the final step, the entire rail is cooled to ambient temperature with a balanced thermal

deformation. This method significantly reduces the final mechanical distortion compensation due to dif-

ferential heating. Wang et al. 55 employed moire interferometry to measure the residual stresses in rails.Precise measurement of residual stresses in the railway tracks aids in estimating the mechanical properties,

including its fatigue life.

Summary

Pearlitic rail steel is extensively used for railway track due to its excellent mechanical properties and wear

resistance. Precise control of chemistry and thermomechanical processing is required to obtain the desired

pearlitic microstructure with fine lamellae spacing. This is achieved by carrying out the pearlitic transfor-

mation at the lowest possible temperature. It is important to precisely control the phase transformation

during thermomechanical processing to obtain the desired microstructure and mechanical properties of the

rail steel. The various physical metallurgy aspects during the production and design of rail steel have beendiscussed.

FIG. 26(a) The residual stress of a rail across a cross section in longitudinal direction at a distance

z of 1500 mm from the edge. (b) The residual stress distribution along the symmetry line (adapted fromRef [53]).



24/26

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