INTRODUCTION AND LITERATURE SURVEYshodhganga.inflibnet.ac.in/bitstream/10603/37110/3/chapter 1.pdfwhite irons are used. Alloyed white irons including abrasion resistant irons find

Introduction 1

INTRODUCTION AND LITERATURE SURVEY

1.1 General

Electric power generation plays a pivotal role in the economic growth of any

country. Power generation is placed in the core sector in our country along

with other areas like space, atomic energy, cement, steel, and agriculture. The

present installed generation capacity in India has crossed 100,000 MW of

power in which 60 % is coal based, 30 % from hydel and the remaining is

from other sources such as nuclear and non conventional resources like wind,

bio-mass etc. [1]. Keeping in view of the short fall of about 12 %, the plan for

the next ten years envisages doubling the capacity with 50 % share coming

from thermal power generation [1, 2]. Presently, the coal consumption for

power production in India is about 220 million tonnes per annum and it will go

up to 420 million tones per annum in the next 10 years [1].

The main source of thermal power generation is coal mineral matter

[3]. The coal available in the country as a fuel for thermal power generation is

of inferior quality owing to high ash content (of about 40 – 50 %) in coal [3,4].

Besides this, it is important to note that coal contains about 10 – 15 % angular

quartz and pyrite (~ 2 %), which are chiefly responsible for wear and erosion

damage of power plant components [3,4] leading to shut down of the plant.

This situation leads to an enormous amount of revenue loss [5] due to the

down time of the system besides disruption of power production and

distribution throughout the grid network.

To combat such wear out of components in power plants, wear

resistant materials are required to be used [4,5]. Such developments in

designing newer materials incidentally benefit other engineering applications

too [6]. Generally, the materials employed for such engineering applications

are carbon steels, low alloy steels, alloy cast irons, manganese steels of

2 Introduction

different grades [7,8]. Alloy cast irons ranging from gray irons to alloyed

white irons are used. Alloyed white irons including abrasion resistant irons

find notable application in a number of engineering industries for wear

resistant applications [9,10]. Among these, nihard irons (Ni-Cr alloyed iron)

and high chromium irons (11 – 30 % Cr) find extensive use for wear resistance

applications [11,12]. During the 70’s, nihard cast irons developed by the

International Nickel Company, USA came into prominence. They contain 3 to

5 % nickel and 5 to 8 % chromium. The well known grades in this category

are nihard II and nihard IV [7,9]. As is known, nihard II features M3C carbides

(M denoting Cr / Fe), whereas nihard IV contains M7C3 type harder carbides

[8,9]. The nihard castings are generally regarded as reliable wear resistant

materials yielding higher life, compared to the traditional engineering steels

like carbon steels. In particular, the nihard family finds application [13] in coal

pulverizers. Nihard II castings find application in bimetallic pulverizer rolls

[6] (Rowland – mills), whereas for other wear resistance situations, nihard IV

is used [13]. Typically the nihard components such as rolls, multiple port

outlets and orifices exhibit a service life of about 2000 to 6000 hours [14]. In

other components, on the other hand, the useful life recorded is sometimes as

low as 1000 hours [14]. To extend this operational life of components to a

value of about 10,000 – 12,000 hours [4,14], all possible efforts and attempts

have been made and these efforts continue to engage the attention of materials

engineers.

1.2 Literature study

As a continuation effort and also as an improvement over nihard, high

chromium irons were introduced into the market during the 80’s [7,14].

Although, the use of high chromium iron castings has increased in the last

decade, it is, however, restricted to proprietary compositions. High chromium

iron features hard discontinuous chromium eutectic carbides (M7C3) in a

martensitic matrix [15,8]. These chromium carbides impart wear as well as

Introduction 3

corrosion resistance [15,8]. But, they are quite brittle and hence do not

withstand shock or impact situations [16,8]. The well known applications of

high chromium irons are in the areas of coal pulverizer/mill parts, coal & ash

handling wear parts, impeller blades for shot blasting equipment, sand &

chemical pump parts, crusher parts for mineral handling, anti-wear plate for

cement clinker cooler, to name a few [6,13]. The high chromium iron was

introduced as hicrome spare parts for ball and race mills by various

manufacturers [14]. The anticipated life could not be obtained since the

required material hardness and the desired microstructural features could not

be achieved. Further, the method of obtaining better wear resistance properties

has been sporadically reported for higher carbon content systems (yielding

higher carbide % in the resulting microstructure) [17], higher cooling rate (fine

dispersion of carbides)[18] and modification by heat treatment [19]. These

diverse attempts have given an indication of achieving better properties

through control of the carbon level. It is known that, as the carbon content is

increased in chromium iron systems, the wear resistance also increases due to

an increase in carbide volume [18]. However, there is a certain upper limit

beyond which an increase in the carbon and chromium contents leads to

development of cracks in the resulting material [20].

To this end, one of the novel methods attempted by several researchers

[21,22,23] has been to alter the microstructure through micro alloying

additions in chromium-rich irons. The work reported by Gundlach and Parks

[24], on the effect of microalloying additions on the abrasive wear, showed

that such additions involving nickel, copper, manganese promote the wear

resistance both in the as-cast and heat-treated conditions. The beneficial effect

of micro alloying additions to high chromium iron system was also reported

by Pearce [6]. He reported that the molybdenum addition increased the

hardenability. Also, controlling the silicon content (to < 1.2 %) suppressed the

pearlite formation, while addition of vanadium increased the toughness. Thus,

4 Introduction

these efforts at micro alloying and the structure-property correlation studies

showed encouraging results.

Another possible method to improve the wear behaviour in chromium

iron system is by heat treatment [25,26,27] which produces a hard martensitic

phase upon fast cooling. Further, the process can be programmed to bring

about a change in the carbide morphological features such as refinement in the

carbide size etc. Laird II [28] reported that control of eutectic carbides through

heat treatment and increasing carbide volume % in the matrix significantly

contributed to improved wear and mechanical properties. The beneficial effect

of higher austenite retention in 28 % chromium bearing white irons was

reported by Lin and Qingde [29] for the conditions involving abrasion and

corrosion-abrasion using a pin on disc abrasion set up. In this investigation, a

fully martensitic alloy obtained through heat treatment showed higher abrasion

resistance using silicon carbide as abrasive disc compared to the heat-treated

austenite rich alloy. The use of soft abrasive disc i.e., garnet, on the other

hand, resulted in higher resistance to abrasion for the heat-treated austenite

rich alloy compared to the as-cast alloy. The researchers attributed this type of

behaviour to higher work hardening nature of the austenite rich alloy [29]. It is

quite clear from the above studies that the desired properties in high chromium

iron system could be achieved by modifying the matrix or by refining the

carbides through heat treatments. The role of heat treatment is, thus, very

clearly brought out by these investigations.

While the formation of martensitic phase can be promoted both in the

as-cast as well as in the heat-treated conditions, a good amount of retained

austenite (RA) is always present in the matrix. This is considered beneficial

from the viewpoint of work hardening ability [12,28], higher crack

propagation resistance and dynamic fracture toughness [17,26]. Work by Fan

et al. [30] on the abrasion resistance and impact fatigue behaviour of high

Introduction 5

chromium iron samples, reports that low carbon martensitic samples possessed

higher crack propagation resistance compared to high carbon martensitic alloy,

as the force required to break the carbide matrix interface was more in the

former. Durman’s [31] investigation on high chromium austenitic iron showed

that depending up on the carbon content, the fracture paths followed different

routes. In chromium irons having lower carbon content, the fracture path

followed the strain induced transformation, whereas at higher levels, the

eutectic carbides controlled the fracture. The work by Xi et al. [32] on the

impact abrasion resistance of high chromium iron showed that at low impact

energy levels, the wear was independent of the retained austenite (RA)

content. On the other hand, at higher impact loads, the wear rate increased

with increase in retained austenite content [32]. In such systems, the RA

content was found to inhibit the initiation and propagation of fatigue cracks

[32], a key factor in any engineering application.

The high chromium iron invariably contains RA % higher than the

desired level, generally, about 10 %. In certain instances, the RA % has been

reported to be as high as 40 % [12]. This was attributed to the lack of close

control in the heat treatment process. The retained austenite can be lowered to

< 5 % in such alloy systems by adopting controlled cryogenic treatment [33].

In this process, the samples were immersed in liquid nitrogen in a controlled

manner, which resulted in transformation from austenite to martensite. This

further increased the wear resistance [33]. The work carried out by Norman et

al. [34] on the abrasion resistance of martensitic white irons reported that the

chilled cast irons exhibited finer carbide structure, pearlite suppression and

superior mechanical properties compared to the sand cast ones due to higher

cooling rate employed in the former.

The third approach to promote the wear resistance in high chromium

iron is by employing higher cooling rate. This is made possible either by

6 Introduction

changing the type of mould from sand to metal and or by providing chilling

arrangement. The use of metal mould or chill promotes faster cooling rate due

to higher thermal conductivity prevalent in the metal mould or chill compared

to sand mould resulting in desirable microstructural features. Also, the higher

cooling rate was reported to provide other features such as good surface finish

[20], less environmental pollution and better dimensional stability in addition

to improved wear resistance not only in the high chromium iron system, but

also in other systems [35].

Now, coming to the applications of high chromium irons under high

stress / gouging conditions such as grinding and crushing operations, they,

besides withstanding wear, should also bear the dynamic stresses [36]. This

obviously poses a problem of finding an ideal compromise between the two

properties, namely, the wear resistance and the impact toughness. In case of

fracture, not only the material toughness matters, but also the complexities

involved like the geometry, distribution of internal stresses, stress

concentration factor, crack formation and propagation have a bearing on the

properties [26]. Hence, the fracture toughness is dependent on several

mechanical, physical and metallurgical parameters.

In order to achieve improved toughness characteristics coupled with

better wear resistance in chromium iron system, several attempts have been

made to alter the matrix for higher retention of austenite in the matrix by

adding elements such as nickel, manganese, copper [8,11]. As nickel and

copper are quite expensive, other alternate materials need to be tried. One such

element is manganese and cost wise cheaper [21,11]. Manganese additions

have been shown to improve the toughness value both at ambient and sub zero

temperatures, by refining morphology of carbides [37]. The usefulness of

manganese addition is to enhance the hardenability independent of the carbon

content [37]. Further, the matrix toughness characteristics improve, as

Introduction 7

manganese is known to be a good austenite stabilizer [8,11,37]. It is also

reported that the manganese addition promotes graphitization tendency [38].

The work carried out by Basak et al. [39] reported that the impact

property was enhanced in sand cooled high chromium iron having manganese

addition up to about 4.4 %, but the improvement seen in respect of the wear

resistance was marginally different. The use of manganese in the range 1 to

4.4 % in chromium iron system and the resulting improved impact behaviour

[39] formed a key point for the initiation of the present investigation.

Stefanescu et al. [40] studied on the structure-property relation in high

chromium (~14 %) cast iron with either manganese or vanadium as alloying

element. While manganese addition from 2 to 4 % is reported to bring down

the abrasion resistance due to the coarseness of the matrix structure, an

increase in vanadium content of the same range, on the other hand, resulted in

the refinement of the matrix structure and thus increasing the abrasion

resistance. The work carried out by Maratray [38] on chromium manganese

alloy systems containing 8 to 14 % chromium, showed that the toughness is

improved with increase in manganese content from 2 to 4 %. The work carried

out by Bolkhovitina et al. [41] on manganese-vanadium irons containing 15 to

30 % manganese reported that these irons possessed good toughness property.

Thus, from the above reported investigations [37,39], the importance of

inclusion of manganese in chromium rich irons is re-emphasized.

From the above literature study, it is understood that the wear resistant

high chromium iron occupies an important place in the ferrous-based systems.

Further, the literature reports reiterate that the wear damage and mechanical

properties of such systems are dictated by process variables such as

composition, cooling rate, heat treatment etc., through microstructural

changes. Therefore, the macroscopic properties have a strong bearing on the

microstructure of the system under study and characterization of the

8 Introduction

microstructural defects on the surface, sub-surface as well as bulk plays a vital

role in understanding structure-property relations in particular the wear

process. From this point of view, the defect characterization and quantification

in terms of its size, concentration and migration is important which can be

assessed using various NDT methods such as acoustic emission, positron

annihilation, low frequency electromagnetic, X- ray based method techniques

etc. Among them, one of the advanced and sensitive methods namely, Positron

Lifetime Spectroscopic (PLS) method seems to give good account of flaws,

defects, porosities, cracks etc in materials [42,43,44], since it has been

established as a powerful and useful tool especially sensitive to small open-

volume defects such as vacancies and small vacancy clusters. Limited

information is available in the literature regarding the defect characterization

in metals especially in steels, wherein in one part, the fatigue damage

accumulation in nickel prior to crack initiation [45] and fatigue damage

detection concerning the extent of damage in steels [46] have been studied and

correlated with PLS parameters. In the other part, how the annihilation of the

defects induced due to radiation in the ferrous alloys affect the behaviour in

terms of migration and annihilation of defect clusters [47] and the surface and

near surface defects formed due to corrosion using slow positron beam

technique [48] have been reported. But the PLS technique as adapted to

characterize chromium manganese iron bulk system does not seem to have

been reported.

Although, high chromium iron shows good promise for wear resistance

applications in thermal power plants, they fail to resist the load under impact

conditions. To supplement the above aspects, some sporadic efforts have been

made to improve the impact behaviour coupled with wear resistance property

through the introduction of manganese to chromium iron, up to 4.4 % [39].

But the addition of manganese in such systems above 4.4 % and damage or

defects characterization by PLS technique has not been reported so far. Hence,

Introduction 9

the aspect of introducing manganese at higher levels (5 and 10 % to chromium

iron) has been taken up in this work as the first objective. The second

objective planned in this work is the influence of cooling rate obtained through

the adoption of metal and sand moulds in chromium manganese iron system,

as other investigators have not looked into this aspect. The next objective,

namely, the effect of casting section size on the wear, mechanical and

metallurgical parameters is looked into as castings having different sizes are

used in engineering industries. Further, the data generated on this aspect will

be very useful to engineering industry. The last objective i.e., heat treatment

effect on the above listed mechanical parameters is looked in to at greater

depth as any study in this field of research will not be complete without this.

To achieve the above cited objectives, the author has used both optical and

scanning electron microscopy for structural examination combined with

positron lifetime spectroscopy for wear damage characterization to find the

correlation between mechanical properties and microstructure.

The literature work emphasized the methods adopted to improve the

wear resistance and impact property in high chromium irons and it has

provided some direction for widening the scope of work further. The key

mechanical property that is looked into is ‘wear’ and what follows is the

coverage on the wear aspects in greater detail.

1.2.1 Wear

Wear is described as the progressive loss of material from the operating

surface due to the relative motion between that surface and the contacting

surface known often by the term counter surface [49]. Wear of metal occurs by

the plastic displacement of the surface and by detachment of particles, which

form wear debris [49]. In metals, this process may occur by contact with other

metals, non-metallic solids, flowing liquids or solid particles or liquid droplets

10 Introduction

entrained in the flow of gases [49]. The wear process may be generally

classified into adhesive, abrasive, erosive, impact, corrosive, fretting and so

on. Of these, adhesive, abrasive and erosive wear phenomena are generally

encountered in engineering applications. As the literature on adhesive wear is

available in abundance and not investigated in the present work, this aspect is

covered in brief. As the emphasis in the present work is laid on abrasion,

erosion and slurry erosion phenomena, the adhesive wear aspects is touched

upon only to form continuity to the related matter in the sections to follow.

1.2.1.1 Adhesive Wear

Adhesive wear is defined as the process occurring due to sliding or rolling

contact between two solid surfaces leading to material transfer between the

two surfaces or loss from either surface. When two surfaces slide on one

another, their topographic features allow only the contact of asperity peaks as

shown in Figure 1.1 [50]. These contact points or ‘Junctures’ represent the real

area of contact. The wear due to the contact of two surfaces has been shown to

follow an equation by Archard [50], which is expressed as wear loss per unit

sliding distance in a simple form

V/S = (β/3) . (W/3σy) ..(1)

where V̀’ is the wear volume, S̀’ is the sliding distance, ‘W’ is the normal

load, ‘σy’’ is the yield stress or flow stress of the material and ‘β‘ is the term

accounting for the probability of a certain number of junctures wearing per

unit sliding distance. The above equation represents a steady state wear.

However, for all practical purposes three regions of wear can be identified

(Figure 1.2) [50, 51, 52]. Region I represents faster wear during the running in

period, while region II corresponds to a slower and steady state wear and

finally the region III represents the terminal conditions. Under high load

conditions, both Region II and III loose their distinct identity [52].

Introduction 11

Figure 1.1: The real contact area (junctures) and apparent (gross) contact area of two surfaces

Figure 1.2: Variation of sliding wear volume with sliding distance

On the other hand, Region II is prolonged in lubricated systems. The

wear in different regions is influenced by various factors such as load, speed,

oxidation, shape and size of the debris, onset of fatigue and micro cracks [52].

The wear process has been explained in literature from the point of view of

12 Introduction

surface and subsurface damage [53,54], known as delamination theory (Figure

1.3). This delamination approach involves the following steps.

a) The deformation patterns in the form of dislocations and vacancies appear

due to sliding action at the surface and subsurface.

b) The formation of voids at the subsurface layers occurs due to the continued

plastic deformation. They increase further in the presence of inclusions and

large precipitate particles at the surface.

c) The voids coalesce either due to the growth or by shearing action of the

surrounding material around hard particles due to the formation of cracks

parallel to the wearing surface.

d) In continuation of the process, the crack after reaching a critical length due

to shearing action yields a sheet like wear particles / debris.

To account for the probability term in Archard’s law favouring the

fatigue theory of wear, Kimura [55] came out with good experimental

evidences supporting the fatigue mechanism by correlating the wear resistance

with fatigue behaviour. The importance of characterizing both fatigue and

wear for analyzing the damage potential of defects and inclusions in materials

under conditions of wear and fatigue, have also been reported by Kimura [55].

Introduction 13

Figure 1.3: Delamination mechanism of adhesive wear

Generally, under adhesive wear situations in ferrous based materials, a

wear resistant white layer is formed with a fine dispersion of carbides and

oxides. These oxide layers possess good wear resistance and aid in reducing

the wear rates [49,54]. The importance of good lubrication in reducing

adhesive wear rates has been reported and well explained in the literature

[41,53]. As lubricants have a great influence in reducing the wear rate, a right

choice of lubricant for a given application has to be made [49]. In the present

investigation, as emphasis is laid on abrasion and erosion behaviours of

chromium manganese irons, these aspects are covered in detail in the

following sections.

14 Introduction

1.2.1.2 Abrasive Wear

Abrasive wear is defined as the wear due to hard protuberances forced against

and moving along a solid surface. It is reported [56] in the literature that the

factors responsible for abrasive wear are hardness, shape and size of the

abrading material. Abrasive wear is generally classified into two types [25, 52,

57]:

a) Two-body abrasion where a hard rough body plough into a softer body; and

b) Three-body abrasion where a third body (usually hard granular matter)

placed between the sliding surfaces gets crushed and cuts grooves.

These types are shown in Figures 1.4 and 1.5 respectively. The two-

body wear is generally a low stress type of wear with particles being

transported across the surface with little breakdown in particle size of the

abrasive [52, 53]. In three-body wear due to the high stress, the particles are

deliberately reduced in size [52, 53]. For all practical purposes, a relative

factor viz., Relative Wear Resistance (RWR) is normally used [54] and

defined as

RWR = (Linear wear of the standard / Linear wear of the material under test) .. (2) As per a published report [52], it is prescribed that the hardness of the

material for abrasion resistance application should be at least 1.3 times that of

the abrasive particles. The hardness of abrasive minerals and ferrous materials

are given in Table 1.1 [52]. From the engineering point of view, abrasive wear

is classified into three specific types (Figure 1.5) [52, 25] and they are briefly

discussed below.

Introduction 15

Figure 1.4: Two body wear and three body wear

Figure 1.5: Types of abrasive wear

1. Gouging abrasion: This takes place due to heavy plastic deformation of a

surface by hard mineral fragments under heavy pressure or impact, causing

deep surface grooving or gouging and removal of relatively large wear debris

particles. Some examples of gouging wear are seen in dragline bucket, rock

crushing.

2. High stress grinding abrasion: This process results in mineral fragments to

fracture under sufficient contact stresses. A few examples of high stress

grinding abrasion are found in pulverizers, ball mills and augers.

16 Introduction

3. Low stress scratching abrasion: This occurs due to cutting or ploughing of

mineral fragments under contact stresses below their crushing strength. The

examples of low stress scratching abrasion are noticed in coal chutes, pump

impellers and ID fans.

Table 1.1: Hardness of abrasives and second phases

Hardness Hardness Mineral

Knoop HV Material or phase

Knoop HV

Talc 20 - Ferrite 235 70-200

Carbon 35 - Pearlite, unalloyed - 250-320

Gypsum 40 36 Pearlite, alloyed - 300-460

Calcite 130 140 Austenite, 12% Mn 305 170-230

Flourite 175 190 Austenite, low alloy - 250-350

Apatite 335 540 Austenite, high Cr iron - 300-600

Glass 455 500 Martensite 500-800 500-1010

Feldspar 550 600-750 Cementite 1025 840-1100

Magnetite 575 - Chromium carbide (CrC3) 1735 1200-1600

Orthoclase 620 - Molybdenum carbide (Mo2C)

1800 1500

Flint 820 950 Tungsten carbide (WC) 1800 2400

Quartz 840 900-1280 Vanadium carbide (VC) 2660 2800

Topaz 1330 1430 Titanium carbide (TiC) 2470 3200

Garnet 1360 - Niobium carbide (NbC) 1900 2400

Emery 1400 - Boron carbide (B4C) 2800 3700

Corondum 2020 1800 - -

Silicon Carbide

2585 2600 - -

Diamond 7575 10000 - -

The evaluation of erosion resistance under solid and slurry conditions

forms an important part of this investigation. Hence, due importance is given

to the erosion phenomenon covering the definition and the parameters

influencing erosion. These are covered in the sections to follow.

Introduction 17

1.2.1.3 Erosive Wear

Erosive wear is defined as the progressive loss of original material from a

solid surface due to mechanical interaction between that surface and a fluid, a

multi-component fluid, impinging solid or liquid particles [57]. Solid particle

erosion is a complex phenomenon in which the three co-existing phases,

conveying fluid, solid particles and the metallic surface interact in many ways

[58]. An understanding of the kinetics of the process [58] involves the analysis

of the following:

a) Properties of the metallic wall, i.e., flow stress, hardness, work hardening

ability, etc.

b) Properties of the solid particle phase, i.e., hardness, solids burden,

velocity, impact angle and particle size as well as its shape.

c) Properties of the fluid phase, i.e., velocity, density, viscosity and flow

regime.

Several theories have been put forward to explain the process of

erosion. Most of the theories explain the target-particle interactions and the

factors influencing the metal removal [58]. The erosion loss is governed by

various parameters such as particle velocity, impact angle, particle size, shape

and distribution. The influence of these parameters on the erosion behaviour

is detailed below [59].

a) Effect of particle velocity

The erosion volume loss is very much influenced by the velocity of the

particle impinging on a target material and this relationship, shown in Figure

1.6 [60], is expressed as

V = K U ..(3)

18 Introduction

where ‘V’ is the volume of the material removed, ‘U’ is the velocity of the

particle, ‘K’ is erosion constant. The velocity exponent ‘n’ is generally in the

range of 2 to 4 [60].

b) Effect of Impact Angle

The particle impingement angle has a direct influence upon the erosion

behaviour. It is reported by Raask [60], that ductile materials exhibit a peak

erosion loss at shallow angles, whereas at normal angles the erosion loss is

very low. On the other hand, brittle materials show least erosion loss at

shallow angles of impact and highest erosion loss at normal impact angles.

The graphical representation of these is shown in Figure 1.7. The work

reported by other researchers [61, 62] as well as by the author [63] on the

effect of the erosion rate with respect to the impact angle is in agreement with

the theoretical predictions.

c) Effect of particle size and its distribution

As reported [60], the particle sizes below 5 µm do not significantly contribute

to erosion. For the particles in the size range 5 to 100 µm, the erosion loss is

marginal and it increases with increase in the particle size as per the equation

[60].

V = C1 dm ..(4)

where ‘C1’ is a constant, ‘d’ is the particle diameter. The exponent ‘m’ is

found to be in the range 0.4 - 0.7 for various target materials [60]. For the

particle sizes above 1000 µm, erosion loss reaches saturation with no further

increase as seen from Figure 1.8 [60]. Raask [60] has further reported that

particles having wider distribution cause higher amount of erosion from a

target material compared to the one coming from a narrower distribution of

particle size.

Introduction 19

Figure 1.6: Effect of particle velocity on erosion volume loss

Figure 1.7: Variation in the erosion volume loss with impact angle

20 Introduction

d) Effect of particle shape and angularity

The particle shape has a significant influence on the erosion volume loss. The

relationship between particle shape and erosion loss, reproduced from a

published report [60], states

V = C2 (Ψ)p .. (4)

where ‘C2’ is a constant and ‘Ψ’ is the particle asperity number. For the

exponent ‘p’, a value of 1.8 has been reported [64]. As regards, the effect of

particle shape on erosion, rounded particles (beach sand) show lower asperity

number, whereas crushed quartz shows higher asperity number due to higher

jaggedness of the particles [14,60]. Therefore, particles having lower asperity

show least erosion and vice versa [14,60]. The work carried out by Liebhard

and Levy [64] on the effect of erodent particle characteristics on the erosion of

metals shows that angular particles cause higher erosion damage compared to

the particles having spherical shape.

e) Effect of impacting particle hardness

The erosion volume loss is found to increase with the impacting particles

hardness as given by the equation [60]

V = C3 (HP)q . . (5)

where ‘C3’ is a constant and ‘HP’ is the particle hardness. The value reported

[60] for the exponent ‘q’ is 2.3. However, it is reported [60] that an increase in

particle hardness beyond 1000 kg/mm2, brings in only marginal changes in the

erosion loss (Figure 1.9).

f) Effect of particle fragmentation

The particle-target surface interaction leads to fracture and fragmentation of

impacting particles. This process is governed by the propensities of impacting

particles, particle velocity as well as the target material hardness [60]. At

particle velocity below 15 m/s and diameter below 10 µm (at any velocity)

fragmentation of particles does not occur [60]. Particles with high initial

Introduction 21

angularity and hardness become frangible and lead to lower erosion [60].

However, when rounded particles fragment (for example: glass spheres) the

resulting angular particles lead to increased erosion [60].

Having covered the aspect of wear and its effect due to various

parameters such as load, speed, velocity, impact angle, particle characteristics

etc, the factors influencing the wear behaviour of the target material such as

composition, hardness, microstructure and heat treatment are explained in

detail below.

Figure1.8: Effects of particle size on erosion wear loss

Figure 1.9: Effect of particle hardness on erosion wear loss

22 Introduction

1.2.2 Factors affecting wear

Wear resistance is not representing the basic material property such as thermal

conductivity, melting point or density. The wear phenomenon is affected by

various factors including processing parameters. Some of the key factors

influencing the wear rate [65] are given below.

a) Design criteria - Transmission of load, type of motion, degree of

lubrication, temperature and environmental factors.

b) Operating conditions such as speed, contact area, contact pressure and

surface condition.

c) Abrasive characteristics such as hardness, shape, size and their

distribution.

d) Material properties: Composition, hardness, microstructure, work

hardening ability and resistance to corrosion.

Having described the factors affecting the wear, the influence of the key

process parameters such as chemical composition, cooling rate and heat

treatment on wear characteristics is considered, as the literature survey clearly

indicates that they control the wear and mechanical characteristics as well as

the metallurgical parameters to a great extent in the ferrous-based system.

Further, the related information is well documented in the literature on high

chromium irons [8,11]. Thus, some of the notable observations in respect of

composition, hardness, microstructure and heat treatment reported in the

literature are summarized below.

Introduction 23

1.2.2. 1 Chemical Composition

a) Carbon

It is emphasized by Gundlach [8] that the amount of carbides is controlled by

the carbon content. Further, the carbides are quite hard and possess good wear

resistance. However, the carbides get fractured very easily under loads and

impact conditions as they are brittle. It is reported that coarse carbides are

formed when the carbon content exceeds the eutectic limit [8,11]. The

eutectic carbon is reported to possess good abrasion resistance. Figure 1.10

shows the effect of carbon content on the impact resistance and it forms an

important basis in the present investigation of optimising wear resistance and

impact behaviour in chromium manganese iron system.

Figure 1.10: Effect of carbon content on impact strength

The inference that emerges from this graph (Figure 1.10) is that, the

impact resistance decreases with increase in carbon content. The importance

of carbon content governing wear, mechanical and metallurgical properties in

chromium irons has also been brought out by other researchers [17,38,66].

24 Introduction

Maratray and Poulalion [67] have reported that carbon content in chromium

iron can increase the amount of austenite, but it is dependent on alloy content,

temperature, kinetics of reaction etc. Some of the observations made by them

support the literature viewpoints with regard to the effect of carbon content on

structure-property relations listed above.

b) Chromium

It is known that chromium is a strong carbide former [15,20]. The carbide

consists of a continuous network of M3C (FeCr)3 C and M7C3 (FeCr)7 C3 type

of carbides surrounded by dendrites of austenite or its transformation products

[8,11]. The M7C3 carbides are much harder (1400 to 1600 HV) compared to

M3C type of carbides (1060 to 1240 HV). It is also known that a discontinuous

M7C3 type surrounded by austenite or its transformation products is formed,

when the chromium content in iron exceeds 10%. This results in the formation

of predominantly the eutectic carbides. Other reports also favour these

findings [6,25,66]. Further, it is reported that the impact resistance (and

fracture toughness) of higher chromium cast irons depends on the toughness of

the matrix rather than that of the carbides [8,11]. However, this point applies

to the irons having low carbon, but not applicable to hypereutectic irons. This

is because of the reason that hypereutectic irons are relatively brittle. The

addition of different levels of chromium content in cast irons and its influence

on the structure have been reported by Barton [68] and the same is reproduced

in Table 1.2. Depending upon the chromium content, the usefulness of cast

iron is exploited.

Introduction 25

Table 1.2: Effect of chromium on microstructure of cast iron with differing Carbon Levels

Chromium % Microstructure

0 Ferrite and coarse graphite

0.3 Less ferrite, some pearlite and finer graphite

0.6 Fine graphite and pearlite

1 Fine graphite, pearlite and small carbide

3 Disappearance of graphite

5 Needle like carbide

10 to 30 M7C3 replace M3C with fine carbide

Beyond 30 % Massive carbide

c) Manganese

Manganese as reported by Gundlach [8], has been known to be a more

potential austenite stabilizer than nickel to provide increased hardenability.

Also, it suppresses pearlite formation. Fairhurst and Rohrig [69] reported that

the tendency of converting austenite to martensite decreases owing to lowering

of Ms (martensitic start) temperature due to manganese addition in chromium

irons. To derive better benefits such as higher impact withstanding capability

in chromium irons in addition to wear resistance, the use of higher levels of

manganese in the range 1 to 4.4 % have been reported by various researchers

[37,38,39,40].

d) Silicon

It is again summarized from the literature [8,11] that the addition of silicon

improves the fluidity of the melt. It is emphasized here that higher hardness is

achieved by increasing the amount of martensite through the addition of

silicon in the range 1 to 1.5 %. Further, it is documented that an increase in

silicon content yields pearlite formation. The addition of ferrosilicon to cast

iron at a later stage helps in increasing the toughness.

26 Introduction

e) Molybdenum

Molybdenum is a very good hardening agent in high chromium white irons. It

is used in heavy section castings to augment hardenability and prevent pearlite

formation [8, 11]. From a typical microstructure reported [70], molybdenum is

generally distributed between the eutectic carbides and the matrix. Further, the

molybdenum addition in small quantities is sufficient to suppress pearlite

formation particularly when used in combination with other elements such as

copper and when the ratio of chromium to carbon is relatively high. It is

reported by Hebbar and Seshan [26] that higher molybdenum addition

promotes increased fracture toughness in 27 % chromium iron due to the

dissolution of molybdenum in the austenite, whereas in 15 % chromium iron

the fracture toughness decreases due to the precipitation of secondary

carbides.

f) Nickel

It is reported in the literature reports that nickel is an austenite stabilizing

agent and its addition increases the toughness property [71]. Further, it is

reported that if the nickel content exceeds the minimum limit required to

inhibit the pearlite formation, this would result in excess amount of RA in the

matrix and consequently lowers the hardness level [8,11].

g) Copper

Copper is quite similar to nickel in its effect on high chromium cast irons. As

reported [8,11], the hardenability increases with increase in copper content

and therefore the retained austenite content also increases. Copper is a good

substitute to nickel as it helps in reducing the requirement of nickel to

suppress the pearlite transformation. Further, it is reported by Srinivasan et al.

[21] that copper addition to chromium iron having chromium level of about

7.5 % helped copper to preferentially get dissolved in austenite resulting in

Introduction 27

decreasing the stability of carbides. It is reported by Dodd and Parks [70] that,

supplementary additions have been recommended in thick sections to improve

the hardenability.

1.2.2.2 Hardness

Hardness is defined [72] as the ability to resist indentation under load or local

plastic deformation. The indentation size due to the application of load on a

material is a measure of hardness. While the wear resistance of any material

increases with increase in hardness, the toughness, on the other hand,

decreases. The hardness value is affected by the compositional variation and

resulting microstructure, which in turn affects the wear resistance and hence it

is considered important. Khruschov [56] reported the existence of a good

correlation between hardness and relative abrasion resistance for different

materials. It is reported [69] that, hardness can be increased by increasing the

carbon content, which in turn increases the wear resistance. However, the

extent of wear resistance improvement is dependent on the nature of hardening

mechanisms involved in different metals and alloys [8,11]. The various phases

present as well as their volume fractions in the matrix in such systems

especially martensitic phase [73] play a significant role on the hardness and

wear resistance both in the as-cast and heat-treated conditions.

1.2.2.3 Microstructure

The micro-structural features in any material provide details on the type of

inclusions, defects, grain boundaries, matrix structure, precipitation of carbide

phases, micro cracks, voids etc and their influence on the resulting material

properties. Therefore, these have an influence on the hardness and wear

resistance of the materials under investigation [8,11, 52].

28 Introduction

a) Carbides

As reported in the literature [8,11], the chromium carbides in high chromium

irons are quite hard and wear resistant. Also, it is known that the wear

resistance is improved in such systems by increasing the amount of carbides

(i.e., by increasing the carbon content), whereas the matrix toughness

improvement is achieved by reducing the carbon content. Figures 1.11 a, b &

c show the influence of carbon content on the shape and distribution of carbide

in chromium iron alloy system. Gundlach [8] reported that when the carbon

content exceeded the eutectic point, large hexagonal carbide rods were formed

at the hyper-eutectic point (Figure 1.11 c). These primary chromium carbides

precipitated prior to eutectic solidification during casting process. They were

quite deleterious to dynamic conditions of (i.e., impact) loading. In such alloy

systems, the mechanical and metallurgical properties were very much

influenced by the chromium level. If the chromium content was less than 12%,

with eutectic or even slightly hypo-eutectic carbon levels, some of the eutectic

carbide may be in the form of cementite, rather than chromium carbide,

resulting in lower wear resistance. The chromium content in the range 12 to 20

% has been reported [8,11] to posses best wear resistance property in the heat-

treated condition due to the formation of martensitic phase. When the

chromium content exceeded 20 % with carbon content limited to the eutectic

composition, the major part of the carbon was used up in forming chromium

carbides, leaving a low carbon martensite matrix, which consequently

improved the wear resistance.

b) Matrix structure

The literature reported that when the microstructure changed progressively

from ferrite through pearlite through bainite to martensite, the wear resistance

increased [8,11]. Lower wear resistance was observed in ferrite structure due

to lower hardness. Further, it was reported [74] that a hard phase formed in the

Introduction 29

matrix such as martensite was responsible for improved hardness as well as

better wear resistant properties. It may be noted that the presence of higher

amount of retained austenite in such systems, lowered the wear resistance.

Figure 1.11: Microstructure of high chromium white iron compositions (a)Low carbon hypoeutectic (b) Eutectic and (c) High carbon hypereutectic

1.2.2.4 Heat Treatment

Heat treatment is also considered an important parameter, which influences

the structure and properties. Certain specific procedures applicable to high

chromium iron, as found in literature, are the following. Heat treatment is used

to obtain a predominantly martensitic matrix with primary carbides in high

chromium irons. The martensitic structure has been reported to favour higher

hardness and wear resistance [8,11]. This is normally done by quenching and

tempering treatments. Gundlach [8] reported the effect of austenitizing

temperature on hardness and austenite content (Figure 1.12) in high chromium

iron and in that work the hardness value decreased with increase in

30 Introduction

austenitizing temperature. Consequently, the austenite content in it increased.

Further, it is noted from the literature that the austenite formed on

solidification at very high temperatures saturated with carbon, chromium and

other alloying elements was quite stable. With decrease in temperature,

chromium and carbon have been reported to combine and form secondary

carbides, reducing the alloy content and thereby the austenite gets destabilized.

The destabilized austenite may get transformed to pearlite, bainite or

martensite depending upon the cooling rate employed. However, carbide

formation becomes sluggish even with moderate cooling rates with

appreciable amounts of super-saturated austenite retained at room temperature.

The as-cast structure is, therefore, often a mixture of pearlite, martensite and

retained austenite.

Figure 1.12: Influence of austenitizing temperature on hardness and retained austenite in high chromium iron

Introduction 31

Normally, thinner sections are predominantly austenitic, while the heavier

sections are pearlitic. It is difficult to obtain a martensitic structure free from

the pearlite in the as-cast condition in these sections. A fully austenitic

structure can be obtained by adjusting the composition with respect to the

section size and the cooling rate. This is made possible in high chromium iron

system by composition control (γ forming element) and use of thin cast

sections. The martensitic structure, on the other hand, is generally obtained by

employing faster cooling rate during the heat treatment process.

a) Soaking procedure

To obtain a predominant martensitic structure with small amount of retained

austenite, it is recommended to destabilize the structure by soaking the casting

within the temperature range 950o C to 1000o C (austenitization temperature)

[8, 11]. This results in decrease of both the chromium and carbon contents in

the matrix due to secondary carbide precipitation. Further, it is reported that

during cooling, the austenite gets transformed to martensite provided the

cooling rate is fast enough to prevent prior transformation to pearlit.

Following the destabilization treatment, it is reported [8,11] that a

variation in hardenability is observed due to the combining effect of carbon

and chromium to form eutectic and secondary carbides with a small amount of

chromium in the alloy still retained in the matrix. Increasing the chromium

content and keeping the carbon level the same can increase hardenability of

the alloy. On the other hand, the hardenability decreases with increase in

carbon content and keeping the chromium content the same.

b) Effect of section size

The reported information [8,11] on the effect of section size on the structure

and hardness of high chromium iron in relation to chromium to carbon ratio is

shown in Figure 1.13. In heavy sections, the formation of pearlite due to

32 Introduction

inadequate hardenability caused a significant reduction in abrasion resistance,

even before its effect on the hardness became apparent. The mixed structure

of martensite and pearlite may also give rise to internal stresses due to

differential changes associated with pearlite transformation taking place first

in an austenitic and later in a martensitic matrix. The castings having mixed

structure show inferior resistances to wear and fracture.

Figure 1.13: Effect of section size on structure and hardness of high chromium cast iron in-relation to chromium and carbon ratio

c) Quenching temperature

It is reported that the transformation characteristics and the final hardness in

chromium iron system are very much influenced by the quenching temperature

by virtue of its effect on the carbon and chromium contents in austenite [8,11].

It is noted that with increase in quenching temperature, the solubility of carbon

in an austenitic matrix increases. At higher carbon levels, higher hardenability

and higher hardness have been reported due to martensite formation following

quenching. A further increase in carbon content leads to an increase in

retained austenite. Consequently the hardness value starts decreasing. The

temperature, at which maximum hardness is attained, increases with increase

in chromium content of the alloy, since chromium increases the temperature of

the transformation from ferrite to austenite. Re-heating the alloy to 400o C -

Introduction 33

600o C produces some destabilization of the retained austenite and this

transforms to martensite on cooling to room temperature.

d) Stress relieving

The tempering treatment is generally carried out to relieve the internal stresses

in the chromium iron alloy system so that a tempered martensitic structure is

obtained. Low temperature tempering in the range 200 – 250° C has been

recommended in the literature [8,11] in view of substantial improvement in

fracture toughness seen as a result of the tempered martensitic structure

formed during the process. Also the process will bring down the residual stress

level. It is reported that tempering at sufficiently higher temperature (above

500° C) would sufficiently reduce the abrasion resistance and therefore, as far

as possible this should be avoided.

1.2.3 Positron Lifetime Spectroscopy (PLS)

The Positron Lifetime Spectroscopy (PLS) is considered as a novel method for

studying the electronic structure, determining the structure, nature, and

concentration of point and extended defects, and investigating the disrupted

surface layers and surface states in metals, alloys, semiconductors, ionic

crystals, and other substances that have firmly established themselves in the

physics and chemistry of solids [43, 75-78]. In the following sections a brief

introduction to this subject is provided.

1.2.3.1 Principle of PLS

Positron (e+) is the antiparticle of electron (e–). The electromagnetic

interaction between electrons and positrons makes possible annihilation of e+ –

e– pairs in which the total energy of the annihilating pair may be transferred to

quanta of the electromagnetic field (photons). Principal channel of this

reaction is the two-photon annihilation,

34 Introduction

e+ – e– � �1 + � 2 ..(6)

The role of conservation of energy and momentum in the process can be

understood as follows. In the center-of-mass frame of the (e+ e–) pair, the

energies of both the annihilation photons are equal to the rest energy of the

electron (positron), E0 = m0c2, and the two photons are emitted in strictly

opposite directions. In the laboratory frame, in which positron is considered to

be at rest, energies of the two annihilation photons are shifted with respect to

E0 by �E ~E ± cPL/2 and the angle between emission directions of the two

photons differs from � by �� ~ PT/m0c. In these expressions, non-relativistic

approximation is used and symbols PL and PT denote longitudinal and

transversal components of the electron momentum, respectively. The electron

(positron) rest mass is designated as m0 and c is the light velocity. Process (6)

is characterized also with annihilation rate � (positron lifetime � = � –1).

Theoretical treatment reveals [75] that � is proportional to the effective

electron density ne sampled by positron, viz. � = � rc2 cne, where rc stands for

the classical electron radius and c stands for the velocity of light [75].

To get an idea of magnitudes of quantities �, �E and �� for e+ – e–

annihilation in condensed matter, one must insert the realistic estimates of

electron concentration ne and electron momentum p into the above

expressions. Conduction electron densities in metals are typically of order of

10 23 m–3 [76] and the positron lifetime in materials are governed by the

electron density. Higher the electron density, shorter is the lifetime [78]. Core

electron moment in atoms may be taken roughly as h/2�ra, where atomic size

is characterized by Bohr’s radius ra and h is Planck’s constant. Then the

following estimates of �, �E and �� may be obtained: (i) Positron lifetimes (τ)

of the order of 200-400 ps are expected in metals [42,77] (ii) Doppler shifts of

Introduction 35

annihilation photon energies, �E ~ 1 keV. (iii) Angular correlation curves

should exhibit widths of a few mrad.

In homogeneous defect-free media, all positrons annihilate with the

same rate �b which is a characteristic of the given material. Due to the

Coulomb repulsion by the positive-ion cores, positron in a condensed medium

preferably resides in the inter-atomic space. At open-volume defects (mono

vacancies, larger vacancy clusters, dislocations etc.), the potential sensed by

the positron is lowered due to reduction in the Coulomb repulsion. As a result,

a localized positron state at the defect can have a lower energy than the state

of delocalized (free) positron. The transition from the delocalized state to the

localized one is called positron trapping. Positron binding energies Eb to

defects like, e.g., mono vacancies are typically of a few eVs [43]. Thermal de-

trapping is impossible from such deep traps and positron remains trapped until

annihilation. If a positron trap is shallow enough (Eb < 0.1 eV), phonon-

assisted de-trapping occurs. As local electron density at the defect site is

lowered compared to that of the unperturbed regions, lifetime �v of the trapped

positrons is correspondingly longer than �b = �b–1. Positron trapping is

characterized by trapping rate, which is proportional to defect concentration c

in the sample, k =�c. Trapping coefficient �, together with annihilation rate �v

= �v–1, are specific for a given kind of defects.

As a result of positron trapping, additional exponential components

occur in the measured positron lifetime spectra. Their appearance can be

explained by the trapping models which give the rate equations for the

positrons annihilating from the delocalised state and from the localised states.

Examples of such models and their use in the quantitative analysis of PLS data

can be found in the literatures [79 - 81]. Recently, the theory of positron

annihilation in solids and solid surfaces has been reviewed by Puska and

36 Introduction

Niemen [43], and on semiconductors [82] by Hautojarvi [80] which gives a

lucid exposure to the subject of Positron Annihilation.

PLS finds its usefulness in polymeric materials to probe free volume

holes with the idea of understanding visco-elastic properties through free

volume quantification [83]. In many macromolecular systems such as

insulators made of polymers, the positron can form a bound state with a host

electron called Positronium (Ps), which is an analogue of the hydrogen atom

[76]. The binding energy of Ps atom is 6.1 eV and its radius is 0.106 nm. Ps

atom can exist in the singlet (para-positronium: p-Ps) or the triplet (ortho-

positronium: o-Ps) state. While the two-photon self-annihilation of p-Ps is an

allowed process, the two-photon annihilation of o-Ps can proceed only via the

pick-off reaction with an electron of the host. Due to exchange interaction of

its electron (that is of Ps) with environmental electrons, Ps atom is repelled to

an unoccupied space in the material called the free volume.

1.2.3.2 Measurement Techniques in Metals and alloys

PAS involves three experimental techniques which originated from nuclear

physics, namely nuclear spectroscopy. This is why the progress made in PAS

is closely related with the achievements in nuclear experimental methods.

Positron-lifetime Spectroscopic measurements (PAS/PLS):

The principle of the PLS measurements involves the use of 22Na positron

source wherein the positron is implanted into the sample almost

simultaneously with the birth � -ray of 1278 keV energy. Lifetimes of

individual positrons (τ) can be measured as time differences between emission

of the birth of � photon and one of the annihilated � photons.

Introduction 37

Doppler broadening of annihilation line (PAS/DB):

It can be measured with a standard �-ray spectrometer equipped with the

HPGe detector. Energy resolution of such devices �E = 1.2 keV at 511 keV is

now being normally achieved.

Angular Correlation of Annihilation Photons (PAS/AC):

In the angular correlation of annihilation radiation, the two annihilation

photons are emitted simultaneously. Thus �� as a function of the transverse

electron momentum component can be measured in coincidence arrangement

with position-sensitive detectors.

1.3 Aim and Scope of the Work

The aim of the present work is to produce wear as well as impact resistant

materials in a ferrous-based system. Chromium, the first material to go with

iron is the natural choice as its hard carbides are expected to improve the wear

resistance. As impact is also a key part of the aim, a γ phase stabilizing

element in the form of manganese is chosen as the second major alloying

element. The manganese level is selected initially at 5 % to go with 16 – 18 %

chromium. The beneficial (or otherwise) effect of raising the level of

manganese to 10 % is next looked into. To enlarge the scope of the present

work further, melting and pouring of such compositions are made in sand and

metal moulds. To further add to the breadth, the section sizes are deliberately

varied in metal mould route. Finally, for better understanding of the processes

listed above (viz., varying manganese level, type of mould adopted and

changing section size) following heat treatment, the wear characterization

involving microstructural examination using optical and scanning electron

microscopy have been carried out. For all the four above mentioned categories

of samples, PLS measurements have been done to study the defect

morphologies and their relationship with wear properties in such systems.

38 Introduction

To achieve the above aim and the broad scope of the work, the review of

literature on the high chromium iron is carefully scrutinized. As emphasized

earlier in the chapter, the thermal power generation in India contributes to the

bulk production of electric power (to the tune of about 60 %) and the materials

employed in these situations should not only possess higher wear resistance

but also better fracture toughness property. If a conventional material like

plain carbon steel is used in such applications, its span of life is quite low.

This may lead to forced plant outage and consequently hampers the power

production leading to loss in revenue to the exchequer. Hence, one has to

resort to the use of good wear resistant materials like high chromium iron as it

has got good potential for wear resistant applications especially in thermal

power plants. These aspects have also been stressed in detail in the earlier

section of this chapter.

It is seen that although wear resistant materials involving chromium

can be made, when it comes to the impact resistance property, especially in

thermal plants, these materials are ineffective. They fail to withstand load as

they are brittle due to the presence of hard carbides and hence get fractured

easily under impact load. This calls for improved impact properties without

significantly sacrificing the wear characteristics. Hence, when the aspect of

having wear and impact properties is looked into, it is seen that very few

investigations aims at solving these twin issues. The scanty reports indicate

that γ stabilizing elements like copper, nickel and manganese can be tried. As

the first two are quite expensive and resources in this country with respect to

their availability are meager, only manganese in chromium is tried in the

present work. To substantiate the selection of manganese, an aspect of using it

from 1 – 4.4 % is reported in one of the literatures reported [39]. Further, there

are no attempts for an in-depth study in chromium manganese iron systems

having manganese content beyond 4.4 %.

Introduction 39

With this encouraging input, a higher level of 5 % in one case and

doubling the manganese level to 10 % in the other is tried in this work by

producing the castings of section size of 24 mm both in metal and sand

moulds. Thus, the work looks at enhanced levels of manganese in iron-

chromium system. The advantage of inclusion of manganese in such system is

to yield material, which has carbide (chromium) forming (favoring resistance

to wear) with γ forming (due to manganese) favoring enhanced fracture

toughness. Hence, the effect of increase in manganese content on the

tribological (abrasion, erosion and slurry erosion), mechanical (impact energy

and hardness), metallurgical (microstructure) properties involving optical &

scanning electron microscopy are looked in to in the first instance. The PLS

technique in respect of increased Mn content has been looked in to from the

point of defect quantification and to study how the PLS parameters get

correlated with tribological parameters.

With this understanding on the effect of increase in manganese

content, the next parameter of importance considered here is the cooling rate.

It is known that metallic structures in general and ferrous-based materials in

particular are quite sensitive to the cooling rates and therefore advantages, if

any, derivable by adoption of varied cooling rates need to be studied. The

point to be noted here is that the work reported in literature on the effect of

cooling rate in chromium manganese iron systems pertains to sand mould

cases only. An investigation on the use of metal mould in such systems has not

been reported so for. Also, there is no information reported on these systems

by using any NDT methods used in such types of studies. Further, the PLS

technique for characterizing this class of materials has never been exploited.

The structural changes that are expected due to the improved heat transfer

characteristics from the metal mould systems and the attendant property

changes as well as defect morphological changes particularly from the wear

and impact angles need to be characterized. Hence, this aspect has been taken

40 Introduction

up in depth in this work for the first time by employing metal moulds for

casting besides the regular sand moulds.

The next factor looked into on a limited scale is the effect of casting

size on the mechanical properties including tribological and metallurgical

properties as well as PLS parameters. Further, how the PLS parameters

affecting the tribological characteristics needs to be studied. It is known that

castings having different section sizes are used in industries. Hence, any study

on section size would benefit the industries. Also, the structure gets affected

due to changes in section size, which would result in attendant property

changes. The defect characteristics also changes on account of varied section

thicknesses. This effect is felt more in the case of metal mould than in sand

type, because of the fact that the metal mould has higher heat dissipation rate

compared to sand mould and hence faster cooling rate prevails in the former.

Thus, in the present work, keeping the metal mould section size at 24 mm as

the reference, the studies are made for the size reduced to 12 mm in one case

and increased to 40 mm in the other case for the 5 % manganese-bearing

samples. However, for the 10 % manganese-bearing case, casting section size

of 24 mm and reducing it to 12 mm section size are only examined.

No mechanical or metallurgical properties evaluation (including

tribological properties) is complete without a stress on heat treatment. It is

well known, from the literature, that heat treatment processes in materials in

general and ferrous based system in particular bring about microstructural

changes such as particle/grain refinement, changes in matrix

structure/appearance, precipitation of secondary phases, change in defect

characteristics etc. Consequently, property changes occur in the system. In the

present case too, the heat treatment cycle (of hardening at austenitizing

temperature of 950° C followed by air quench and tempered at 200° C) is

evolved based on the literature study and the same is given to the select alloy

Introduction 41

system, viz., chromium manganese iron. It is worth noting here that in these

systems, the other researchers have not looked at this aspect before. Therefore,

the above heat treatment schedule is given to all the three categories of

samples listed earlier and the resulting microstructure and properties changes

have been studied using various techniques outlined above including PLS for

achieving the aim outlined earlier.

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42 Introduction

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