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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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