4-Characterisation and machinability of high chromium

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Wear 460-461 (2020) 203463
Available online 28 August 2020
0043-1648/© 2020 Elsevier B.V. All rights reserved.
Characterisation and machinability of high chromium hardened white cast 
iron with and without the addition of niobium 
Anderson Edson da Silva a,*, Ismael Nogueira Rabelo de Melo b, Ivete Peixoto Pinheiro a, 
Leonardo Roberto da Silva a 
a Materials Engineering Department, Federal Centre of Technological Education of Minas Gerais, CEP, 30421-169, Belo Horizonte, MG, Brazil 
b Federal Institute of Minas Gerais, Ibirité, CEP, 32407-190, MG, Brazil 
A R T I C L E I N F O 
Keywords: 
Machinability 
High chromium white cast iron (HCWCI) 
Characterisation 
Niobium 
pcBN tools 
Wear mechanisms 
A B S T R A C T 
High chromium white cast iron (HCWCI) is considered a hard machinable material because of its high carbide 
content, which when associated with a matrix, forms a rigid and very hard microstructure. This work evaluates 
the effect of the addition of 0.5% niobium (Nb) on the machinability of a HCWCI alloy with 25.6% chromium and 
3.2% carbon. The alloys were cast, heat-treated and characterised. The characterisation was performed by X-ray 
diffractometry, scanning electron microscopy, hardness and carbide volumetric fraction (CVF). The machin-
ability was analysed in materials with and without the addition of Nb during the turning process using poly-
crystalline cubic boron nitride cutting tools, while varying the cutting speed in dry cutting and lubri-cooling 
conditions. The output parameters analysed were the tool lifetime, roughness and wear mechanisms. The Nb 
addition promoted a decrease of CVF and hardness, also resulted in the microstructure refinement, contributing 
positively to the machinability of HCWCI under both cooling conditions. In general, the use of cutting fluid 
resulted in lower roughness values and improved life performance of cutting tools, influencing the reduction of 
the abrasion wear mechanism. 
1. Introduction 
High chromium hardened white cast iron (HCWCI) is widely used in 
processes having high mechanical requirements and wear resistance but 
having limited toughness. Its wear resistance is directly linked to its 
microstructure and the addition of alloy elements [1]. These materials 
have been frequently utilised for centuries and many studies on their 
metallurgical characteristics and wear behaviour have been conducted 
[2]. Its high resistance to wear is mainly linked to its microstructure 
formed after hardening by a martensitic matrix, with a large occurrence 
of M7C3-type primary and eutectic carbides [3]. 
Bouhamla et al. [4] added niobium (Nb) to an HCWCI containing 
15% chromium (Cr) and 2.31% carbon (C), with Nb content ranging 
from 0.5 to 3.0%. Filipovic et al. [5] utilised similar Nb contents, but the 
HCWCI had approximately 17% Cr and 2.96% C. Correa et al. [6] added 
7% Nb to an HCWCI containing 24.5% Cr and 4.85% C. These works, 
among others, disclosed considerable wear resistance improvements in 
Nb-enriched HCWCI. According to Penagos et al. [7], HCWCI alloys with 
low Nb content can offer good wear resistance results and a good 
cost-benefit ratio for industrial applications. Percentages of addition 
between 0.5% and 1.0% Nb show a good cost benefit in terms of 
increasing the wear resistance of HCWCI. These gains indicate that the 
use of Nb to increase the wear resistance of this material may be feasible 
and therefore may be widely used in the coming decades. 
Oliveira et al. [8] indicated that mechanical components subject to 
wear must exhibit high hardness (greater than 45 Rockwell C hardness 
(HRC)), which is common for finishing operations on these parts in the 
rectification process. However, with the development of ceramic tools 
and polycrystalline cubic boron nitride (pcBN), turning has become a 
viable alternative. Turning of hardened materials is typically used on 
steels with restricted carbide volumetric fraction (CVF), which produces 
challenges for the machining of HCWCI mainly because of its high CVF, 
which can lead to high wear rates and cutting-edge failures. 
HCWCI machining with carbide tools, even of special or coated 
grades, is not efficient. The high hardness of the material requires the 
use of specific tool materials, and pcBN tools have been suggested [9]. 
Turning of high hardness steels is largely conducted using pcBN 
tools, even on high hardness materials (40–60 HRC) obtained after 
* Corresponding author. 
E-mail address: andersonsilva@cefetmg.br (A.E. Silva). 
Contents lists available at ScienceDirect 
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https://doi.org/10.1016/j.wear.2020.203463 
Received 3 April 2020; Received in revised form 8 July 2020; Accepted 25 August 2020 
mailto:andersonsilva@cefetmg.br
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quenching heat treatments. However, mechanical components submit-
ted to severe abrasive loads show, in addition to high matrix hardness, 
the presence of a large volumetric fraction of rigid particles in their 
microstructure [8]. 
Research is lacking on the machinability of HCWCI, even though this 
material is widely used in industries such as in mining. With the 
development of new materials for wear-resistant cutting tools in the 
machining of hardened materials such as pcBN, the machining of high 
hardness materials has been extensively studied in the last few decades. 
In this study, a hardened HCWCI alloy, with and without the addition 
of Nb, was cast and characterised. The characterisation was performed 
by X-ray diffractometry (XRD), scanning electron microscopy (SEM), 
HRC and by determining the CVF. This alloy was submitted to the 
turning process using pcBN tools by varying the cutting speed under dry 
conditions and when using cutting fluid. This study expands research on 
the machinability of high hardness materials with a high volume frac-
tion of carbides with the goal of seeking alternatives in the machining of 
HCWCI in relation to the rectification process. 
2. Experimental procedure 
Samples were manufactured through the casting process using a 
HCWCI ASTM A 532 Class III Type A rotor. The melting process was 
performed in a 300-Kg capacity induction furnace. The pouring tem-
perature was 1550 ◦C ± 50 ◦C. Half of the charge was poured into the 
moulds, and an Fe–Nb alloy with a 21.5 μm mean particle size was added 
to the other half, with 15 min of dissolution time. 
The geometry of the samples was developed to derive greater use of 
the material in turning tests with a finishing procedure during the facing 
operation. Fig. 1 shows the design of the specimens for the machinability 
tests as well as the surface to be machined. 
The samples were annealed at 700 ◦C for 6 h and cooled in an oven. 
After the cooling, they were pre-machined and later passed through 
austenite destabilisation at 1050 ◦C for 2 h with cooling by forced 
ventilation. Finally, they were tempered at 200 ◦C for 2 h, with the aim 
of obtaining a martensitic matrix according to Ortega-Cubillos et al. [10] 
and Tabrett et al. [11]. 
Alloys chemical analyses were performed by optical emission spec-
trometry using a SPECTROMAXx spectrometer. Crystalline structures 
characterisation was determined by XRD. The equipment used was a 
Philips-PANalytical XRD model PW 1710, where the X-ray tube had Cu 
Kα radiation of λ = 1,54,056 Å. The test was performed at a scanning 
interval of from 30 to 100◦ (2θ) at 0.02◦ steps with a time per step of 1 s. 
The voltage used was 50 kV and the current was 35 mA. 
For microstructural analysis, a JEOL SEM model JSM-6510 LV 
equipped with a Thermo Fisher Scientific NSS dispersive energy spec-
trometerwas used. The tools were analysed in a SHIMADZU SSX-550 
SEM to evaluate the wear mechanisms. 
The Fiji software (Fiji is just Image J) was used to check the CVF. The 
samples were polished and attacked with Vilella. Images were generated 
by SEM at 500× magnification. 
The material hardness were measure by a A Dura Visio DV30 uni-
versal durometer. Samples were cut by wire electro erosion in the 
perpendicular direction to the surface to be machined to analyse the 
material hardness as a function of the depth. The hardness was measured 
every 3 mm in depth until reaching 18-mm depth with 3 measurements 
at each depth. 
Experiments were performed on a horizontal lathe with compu-
terised numeric command and operating system FANUC 32i-B, INDEX 
brand, model IT 600. The main characteristics of the machine were: a 
40.3-kW power engine with a 4000-RPM maximum rotation and a tool 
holding tower with 14 positions and a 600-mm distance between tips; a 
hydraulic plate with a passage diameter of 65 mm and a 620-mm 
maximum allowable diameter. The performed tests utilised tools and a 
tool holder from Sandvik Coromant. 
The pcBN tool meets SNGA120408 specification class CB7025 with 
no coating, with 60% cBN and bimodal grain distribution (1 and 3 μm). 
The tool holder utilised was a DSSNR 2020K 12, which resulted in the 
following main angles: main position angle (χr) of 45◦, secondary po-
sition angle (χr) of 45◦, exit angle (γ0) of -6◦ and inclination angle (γs) of 
-6◦. 
To increase the rigidity and stability of the tests, the fixation system 
was changed to allow a greater contact area between the sample and 
jaw. The jaw was modified to obtain a contact surface with a radius like 
that of the part, thus allowing better fixation and guaranteeing external 
concentricity during the tests. The facing direction was from the end to 
the centre. A backrest was also made that limited the positioning of the 
part during the tests. This backrest is important to prevent a possible 
displacement of the sample in the direction of the spindle during the 
experiments and to facilitate the process of referencing the system 
during interruptions for tool wear and roughness measurements. 
After each interruption, when the part in the machine was replaced, 
it was supported and forced against this backrest at the time of clamping, 
thus ensuring the referencing. Fig. 2 shows the clamping system used in 
the machining experiments. 
The tests were developed using the same parameters for materials 
with and without Nb addition. The 0.20-mm cutting depth and 0.08- 
mm/rev feed were kept constant in all experiments. The cutting 
speeds used were 100, 200 and 300 m/min under dry conditions and 
with cutting fluid. Machining times were calculated using Mastercam 
programming software version X7. 
Syntyl 9902 synthetic fluid was used in the cutting fluid tests. The 
concentration used was 7.5% of fluid, volume basis, verified by a 
portable refractometer Atago MASTER-α series, Brix 0.0–33.0%, while 
observing the correction of refraction reading as provided by the 
manufacturer. 
The lubri-cooling system adopted was jet lubrication at an approxi-
mately 50 L/min flow rate, with the primary and secondary lubri- 
cooling system of the equipment activated. The primary and second-
ary lubri-cooling were applied to the tool exit surface and clearance 
surface, respectively, as shown in Fig. 3. 
The flank wear (VBB) was measured in an optical microscope every 
two or four passes depending on the evolution of wear. The roughness of 
the machined surface was determined by a Mitutoyo SJ 310 portable 
roughness meter with a sampling length of 0.8 mm, where eight mea-
surements were taken spaced at 45◦. These values were measured and 
recorded at each set of passes of the tool to analyse the evolution of each Fig. 1. Machinability test specimen. 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
3
of these variables. The measured roughness parameters were Ra and Rt. 
The tests were interrupted when the maximum tool flank wear or cutting 
time of 30 min was reached. Interruptions also occurred when there was 
micro-chipping or damage to the part or tool. These parameters were 
adopted according to the pre-test results to maintain the integrity of the 
cutting edge, enabling a later study of wear mechanisms developed 
during machining. Due high cost of the cutting tools and workpiece 
material, the values shown in all graphs are the average of three tests. 
The methods used for data analysis and statistical treatment were the 
bicaudal t-test when only two populations were to be compared and 
analysis of variance (ANOVA, single factor) was used to compare more 
than two populations. 
3. Results and discussion 
3.1. Chemical analysis 
Table 1 shows a comparison of the chemical compositions of the 
reference alloy (specified by ASTM A 532 standard [12] with those of 
alloy A (0% Nb) and alloy B (0.5% Nb). 
The chemical compositions of all elements were obtained within the 
limits established by ASTM A 532 [12]. Although the two alloys origi-
nated from the same batch, a small change could be perceived between 
alloys A and B, mainly in their C and Cr elements, which showed a small 
decrease in alloy B as compared to alloy A. This decrease could be 
explained by the fact that alloy B was longer cast due to the time 
required for dissolution of Nb and by the fact that the added mass in-
creases the total mass and reduces a portion of the contents of the other 
elements. Following the removal of the material referring to alloy A, 
alloy B remained in a liquid state for an additional time of approximately 
25 min, corresponding to the dissolution of the Nb and the preparations 
for pouring. During this additional time, these metals may have under-
gone burning and oxidation processes. 
3.2. Hardness 
In Fig. 4 (a)–(b), the material hardness behaviour is apparent, as the 
depth increases in the part for both alloys with and without Nb addition. 
Table 2 shows the hardness ANOVA as a function of the depth for 
alloys with and without Nb addition. 
For p > 0.05 and F < F critical, it can be stated that, statistically, no 
significant variation occurs between the hardness. As the part was 40 
mm thick and the hardness was analysed up to 18 mm deep, affirming 
that the hardness remained homogeneous throughout the part during 
the machining experiments was possible. Hardness was monitored 
during the machining tests and no significant variations in hardness 
were observed. 
3.3. Microstructure 
The microstructures of the tempered and quenched alloys are shown 
in Fig. 5. The alloy without the Nb addition (Fig. 5(a)) was hypereutectic 
and remained in this condition even after the addition of Nb (Fig. 5(b)). 
It was also possible to observe that the detached carbide rods were larger 
than 300 μm in the longitudinal position. Chung et al. [13] indicated 
that additions of Nb in certain concentrations can promote changes in 
the microstructure from hypereutectic to eutectic and hypoeutectic, it 
happens because of Nb carbides precipitation witch result in carbon 
impoverishment in the molten material, thereby affecting its 
solidification. 
After the addition of Nb, identifying primary carbides in the alloy 
Fig. 2. Clamping system of the part during machining tests. 
Fig. 3. Utilised primary and secondary lubri-cooling system. 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
4
was still possible, indicating that the alloy remained hypereutectic. The 
formation of Nb carbides depletedthe liquid but not to the point of 
changing its microstructure to eutectic or hypoeutectic. This fact can be 
explained by the low addition of Nb. 
A comparison of Fig. 5(a) and (b) reveals that the microstructure was 
refined after the addition of Nb. According to Filipovic et al. [14], Nb 
carbides are the first to be formed, followed by M7C3 carbides in hy-
pereutectic alloys. As C is the main element for determining the amount 
of carbide in HCWCI, the number of carbides in the alloy decrease as the 
Nb content increases. 
Identifying Nb carbides in small magnifications, as in Fig. 5(b), is 
difficult because in these magnifications the Nb carbide has a colour 
similar to that of the matrix. Another fact that makes visualisation at 
small magnifications difficult is the fact that Nb carbide is generally 
smaller in size than Cr carbide. To show the distribution of Nb carbide in 
the material structure, we conducted a chemical mapping, as shown in 
Fig. 6. 
Fig. 6 shows that carbides were homogenously distributed and sized 
Table 1 
Chemical composition of cast alloys based on weight (%). 
Alloy Class Type C Mn Si Ni Cr Mo Cu Nb 
ASTM A532 III A 2.0–3.3 2.0 max 1.5 max 2.5 max 23.0–30.0 3.0 max 1.2 max – 
A 0% Nb 3.222 0.942 1.230 0.464 25.594 0.090 0.136 0.117 
B 0.5% Nb 3.092 0.931 1.217 0.467 25.313 0.089 0.138 0.610 
Fig. 4. Hardness as a function of depth for alloys with and without Nb addition. 
Table 2 
ANOVA of the hardness of alloys with and without Nb addition as a function of 
depth. 
Condition p value α F F critical Conclusion 
0% Nb 0.262 0.05 1.424 2.772 same 
0.5% Nb 0.468 0.05 0.958 2.772 same 
Fig. 5. Microstructure of tempered and quenched alloys with and without Nb addition. 
A.E. Silva et al. 
astm:A532
Wear 460-461 (2020) 203463
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throughout the material and presented compact morphologies. Nb car-
bides were approximately 5 μm in size, whereas Cr carbides were larger 
than 300 μm in some cases in their longitudinal sections. The observed 
matrix following all phases of treatment was martensitic, with the 
occurrence of secondary carbides. Primary Cr and Nb carbides did not 
undergo morphological changes due to heat treatment. The morphol-
ogies of the microstructures of the alloys are displayed in Fig. 7, with the 
martensitic matrix (M), secondary carbides (SC), Cr primary carbides 
(M7C3) and Nb carbides (NbC) highlighted. 
The chemical mappings shown in Figs. 8 and 9 illustrate the Cr and 
Nb carbides. They also enable the secondary carbides in the martensitic 
matrix to be visualised. These secondary carbides are also of M7C3 type 
but originated during austenite destabilisation. 
Nb carbides appeared in compact form but were found basically in 
two ways (as shown in Fig. 10): associated with a primary Cr carbide 
(represented by the letter A) or isolated on the martensitic matrix 
without any connection with the Cr carbide (represented by the letter B). 
The Nb and Cr carbides association suggested that the former served 
as a substrate for the nucleation of the latter. Fig. 11 shows the contact 
zone between the Cr and Nb carbides. 
Filipovic et al. [14] indicated that Nb carbides are the first to solidify 
in the alloy and can act as a substrate for the nucleation of austenite 
dendrites in hypoeutectic alloys. However, Fig. 11 shows that the Nb 
carbide is mainly associated with the M7C3 carbide. As seen in Fig. 5, the 
alloys solidified under hypereutectic conditions, and thus the solidifi-
cation began with the formation of M7C3 Cr carbides at temperatures of 
approximately 1300 ◦C. The impoverishment of the liquid caused by the 
formation of Cr and C rich carbides produced a displacement of the alloy 
towards the eutectic line, where the simultaneous solidification of car-
bides and austenite occurs. 
Because niobium carbides act as nucleating agents, it is possible that 
nearly all Nb carbides found were associated with Cr carbides. This 
could be explained by the nucleation tendency of the Cr carbides in 
relation to Nb carbides. Thus, Nb carbides that were formed and served 
as nuclei for the formation of Cr carbides that were later structured were 
predominantly associated with Nb carbides. 
3.4. CVF 
Carbides are very hard and abrasive particles that hinder a material’s 
machinability. HCWCI has numerous carbides dispersed in its matrix. 
This results in a high hardness material with difficult machinability, 
which in turn increases cutting forces and causes high wear rates in 
cutting tools [9]. The CVF of alloys with and without Nb addition are 
shown in Fig. 12. 
A t-test of the samples was conducted (Table 3) to evaluate if a sta-
tistical difference existed for the CVF between the alloys based on a 5% 
degree of significance. The values were in fact statistically different. 
The inclusion of Nb in the alloy caused an average CVF reduction of 
approximately 9%, and this reduction agrees with the results of Ibrahim 
et al. [15]. Microstructural changes could be explained by the fact that 
the Nb inserted in the bath appropriated a part of the C to form a Nb 
carbide. Thus, a depletion of the liquid occurred, thereby compromising 
the formation of M7C3 Cr carbides. The result was a microstructure with 
a lower and more refined CVF. 
3.5. X-ray diffraction 
Fig. 13 presents the results of HCWCI alloys X-ray diffraction with 
and without Nb addition under thermally treated conditions after 
hardening and quenching. 
The analysis revealed the presence of four phases: austenite (γ), M7C3 
Cr carbides (C), Nb carbides (N) and martensite/ferrite (M), where the 
peaks may correspond to both ferrite and martensite, as the peaks of 
these two phases were coincident. 
Nb carbide was identified at a 35◦ angle position in the Nb added 
alloy. M7C3 Cr carbides were identified in all alloys, and the peaks 
referring to this phase were identified in positions 39, 43, 44.5, 51 and 
82.5◦. These peaks were found in a similar manner by other authors, 
including Higuera Cobos et al. [16] and Ibrahim et al. [15]. The peaks at 
positions 44.5 and 82.5◦ indicated a coincidence of peaks between Cr 
carbides and ferrite/martensite, just as the peak at position 51◦ indi-
cated an overlap of the austenite and Cr carbides peaks. This overlap was 
similarly identified by Higuera Cobos et al. [16]. 
3.6. Analysis of HCWCI machinability 
3.6.1. Lifetimes of cutting tools 
Fig. 14 shows the results for the lifetimes of the cutting tools when 
turning HCWCI as a function of the following variables: materials with 
Fig. 6. Chemical mapping of tempered and quenched Nb-containing alloys. 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
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and without Nb addition, cutting speed and cooling conditions. The 
cooling conditions were identified as D (dry) for dry tests and C (cooling) 
for tests using cutting fluid. 
The main parameter affecting the wear of the cutting tools when 
machining HCWCI was the cutting speed. As the cutting speed increases, 
the tool lifetime decreases significantly, probably because of the 
increased heat generated by the process combined with the harsh 
machining conditions imposed by the material such as high hardness 
and high CVF. This behaviour was expected. Mia and Dhar [17] reported 
that the cutting temperature is a major variable that promotes wear in 
cutting tools. An increase in cutting speed strongly affects the heat 
generated during the machining process, thus causing a higher energy 
build-up but without increasing the tool area which could absorb this 
heat. Bouachaet al. [18] also indicated that the cutting speed is the 
parameter that most affects wear in pcBN tools when turning high 
hardness materials. Under dry conditions in the alloy with the Nb 
addition and under cooling conditions without Nb addition (both at a 
100 m/min cutting speed), the tools withstood 30 min of testing without 
reaching end-of-life conditions. In all other situations, maximum flank 
wear was reached or the test had to be stopped due to tool damage. 
Fig. 15 highlights the evolution of cutting tool wear as a function of the 
volume removed. 
Under dry cutting conditions at a 200 m/min cutting speed (as 
compared to 100 m/min), with other cutting parameters retained, the 
reduction in tool lifetime was approximately 78% for materials both 
with and without Nb addition. At 300 m/min speed, the reductions in 
tool lifetimes were greater, reaching approximately 92% in both alloys 
as compared to a 100-m/min cutting speed. When 200 and 300 m/min 
cutting speeds were compared, where the increase in cutting speed was 
50%, an even greater accelerated wear could be observed, with a 
decrease in tool life of approximately 65%. This behaviour indicated 
that the increase in cutting speed produced a considerable increase in 
machining temperature, and this temperature was directly associated 
with the wear mechanisms, which accelerate exponentially with an in-
crease in temperature, thereby accelerating the end of tool lifetime. In 
general, the cutting fluid increased tool life when machining HCWCI. 
These gains were 20% in terms of tool lifetime in material with Nb 
addition at a 300 m/min cutting speed, whereas the gain in material 
without Nb addition under the same conditions was 60%. A similar yield 
was found at a 100 m/min cutting speed in material without Nb 
Fig. 7. Morphologies of the microstructures of tempered and quenched alloys. 
Fig. 8. Chemical mappings of tempered and quenched alloys without the addition of Nb. 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
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addition, where a 50% gain was achieved for dry cutting. However, in 
this case, the test was interrupted because the tool reached 30 min of 
machining, which indicated that the result would have been higher if the 
tool had continued machining until it reached its end of life due to wear. 
The most significant performance was at a 200 m/min cutting speed, 
when material with Nb addition had a gain of approximately 162% of its 
life under dry conditions, whereas material without Nb addition had a 
gain of 209% of life when using cutting fluid. The only negative result 
was at a 100 m/min cutting speed in the Nb added material, when 
micro-chipping (possibly caused by increased machining forces) forced 
the tests to be stopped, thereby producing a tool life that was shorter 
than under dry conditions. As observed afterwards, the wear 
mechanisms changed dramatically with the use of cutting fluid, and 
these mechanisms directly affect cutting tool lifetime. One of the major 
factors that potentially affects tool wear is the cutting temperature. The 
temperature can be considerably changed when cutting fluid is 
employed. The smaller gains obtained under the 300 m/min cutting 
speed condition could be explained by the low efficiency of the lubri- 
cooling. In other words, at a higher cutting speed, the fluid jet may 
not have efficiently reached regions close to the cutting areas. Therefore, 
its contribution to the final performance becomes less significant. Under 
the 100 m/min cutting speed condition, despite the material without Nb 
addition having responded well to the use of the fluid, the material with 
the Nb addition showed a lower response than under the dry conditions. 
Fig. 9. Chemical mappings of tempered and quenched Nb-containing alloys. 
Fig. 10. Nb carbide. 
A.E. Silva et al. 
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At lower cutting speeds, the cutting fluid penetration efficiency 
increased. However, under these conditions, lubri-cooling may not yield 
as many benefits to HCWCI machining. The best results under the lubri- 
cooling conditions were found at the 200 m/min cutting speed, where 
the highest volume of removed material from among the conditions 
analysed was achieved. This may have contributed to a more efficient 
working condition. The cutting fluid was probably able to contribute 
more effectively, enabling areas near the cut to be reached, providing 
the necessary lubrication and cooling to the process, and without 
excessively reducing the cutting temperature. 
3.6.2. Effect of Nb addition 
The addition of Nb to HCWCI could initially be expected to be an 
aggravating factor in the machinability of the material. However, test 
results revealed that this addition considerably improved the machin-
ability of the material, particularly under the dry cutting conditions. In 
general, the tool life increased by approximately 50% as compared to 
the material without Nb addition. At a 300 m/min speed, the increase in 
tool lifetime in the Nb-added material reached 100%. The addition of Nb 
to the alloy caused the formation of Nb carbides. Nb carbides are the first 
to form and remove C from the bath, making the rest of the alloy poorer 
in C. This C deficiency causes the alloy to have nearly a eutectic 
composition, thereby reducing the average size and volume of primary 
carbides. The reduction of the primary carbide fraction results in lower 
cutting resistance and contributes significantly to a decrease in tool 
wear. Unlike most steels that have a virtually homogeneous micro-
structure, the microstructure of heat-treated HCWCI consists of a pre-
dominant matrix of martensite with precipitated secondary carbides. 
This matrix surrounds the eutectic and primary carbides, thus forming a 
heterogeneous microstructure, where the carbides are harder than the 
matrix. The heterogeneity of the alloy microstructure produces insta-
bility during cutting as a function of the machining variation of the 
microstructure. In our study, we determined that the change in carbide 
morphology, mainly derived from a decrease their average size and 
associated with the carbide volumetric reduction, may have contributed 
significantly to an increase in tool life. The average hardness of the 
material without Nb addition was 61.7 HRC and dropped to 60.2 HRC in 
the material with Nb addition. However, this reduction, which was 
approximately 2.5%, would by itself not fully explain the better 
machining performance in the Nb added alloy. Ånmark et al. [19] and 
Dosbaeva et al. [20] described that macro hardness is insufficient to 
determine the machinability of a material, particularly in materials with 
a high abrasive index. Chen et al. [21] indicated that the abrasive index 
of a material is directly related to the CVF and hardness of the matrix. 
This improvement in machinability was most likely achieved through 
characteristics conferred on the material by the addition of Nb. The main 
characteristics that may be associated with this gain are: 2.5% decrease 
in hardness, refinement of the microstructure, 9% reduction in volu-
metric fraction of carbides, and change in the morphology of M7C3 Cr 
carbides. 
Fig. 11. Association of Cr and Nb carbides. 
Fig. 12. CVF in alloys with and without Nb addition. 
Table 3 
CVF t-test. 
Analysis p value α Conclusion 
CVF 0.00617 0.05 Different 
A.E. Silva et al.Wear 460-461 (2020) 203463
9
3.6.3. Roughness analysis 
Part roughness is essential to the analysis of surface finishing quality 
and can inclusively be used to assess the end of the tool life. The 
following analyses considered the roughness evolution in terms of both 
the volume of removed material and the flank wear of the cutting tools. 
Mean arithmetic deviation (Ra). 
Fig. 16 allows the mean arithmetic deviation (Ra) to be analysed at 
the ends of tests under all conditions. 
Fig. 16 shows that the cutting speed had practically no significant 
influence on the roughness in HCWCI machining. The alloy appears to 
have influenced the roughness, as alloys without Nb addition had higher 
roughness values than those with Nb addition. Most notable is the 
reduction in roughness due to the use of cutting fluid, where a significant 
decrease in roughness can be seen in both alloys and at the indicated 
cutting speeds. 
Fig. 17 demonstrates the mean arithmetic deviation (Ra) as a func-
tion of the volume of removed material, where the roughness can be 
observed as the volume of removed material. As a result, the wear of the 
tool increases. 
Both materials displayed a mean roughness of less than 0.40 μm in 
the initial passes. The exception was at the 300 m/min cutting speed 
during dry cutting, when slightly higher values were found in the initial 
passes. This does not mean that the cutting speed affected the roughness 
values but that in these situations, at the intervals established for 
Fig. 13. XRD of alloys with and without Nb addition after hardening and quenching thermal treatments. 
Fig. 14. Lifetimes of cutting tools. 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
10
Fig. 15. Evolution of flank wear. 
Fig. 16. Mean arithmetic deviation (Ra). 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
11
measurement, the tools already presented higher flank wear values 
during the first passes, thus directly influencing the part surface quality. 
Nb added alloys showed lower average roughness values than alloys 
without Nb addition during machining. The carbides contained in the 
material matrix could be broken or removed from the matrix by the 
action of the cutting tool. Analysing the chips generated during HCWCI 
machining, Oliveira et al. [22] noted that the action of the tool caused 
the M7C3 primary Cr carbides to break. After being segmented, these 
carbides could be rearranged in the microstructure of the material, both 
in the chip and on the part surface. Segmented carbides rearranged on 
the surface of the part could influence its surface properties, including 
roughness. As previously mentioned, M7C3 Cr primary carbide rods in 
the longitudinal direction can reach sizes greater than 300 μm in both 
alloys. In finishing machining, these carbides may be larger than the 
depth of the cut itself, reinforcing the possibility of fragmentation of 
these carbides during machining. 
Considering M7C3 Cr primary carbides fragmentation, as the alloy 
without Nb addition presented higher carbides as compared to the Nb 
added alloy, whereas the alloy with Nb addition presented a more 
refined microstructure, the carbides of the alloy without Nb addition 
when fragmented may have interfered more significantly in the surface 
quality of the part. This would justify the higher roughness, even in 
similar values of flank wear in cutting tools. In addition, the alloys 
without Nb addition presented a higher volumetric fraction of carbides, 
thus increasing the surface area of fragmented carbides in the same 
sample length. This thus contributed significantly to the roughness of 
the machined parts. 
The evolution of roughness in dry cutting increased and was prac-
tically linear in all tests. Very similar results were generated, regardless 
of the cutting speed. The roughness appeared to be directly linked to tool 
wear and had negligible effect on the cutting speed. However, higher 
cutting speeds generally cause higher tool wear, and thus roughness 
increases more sharply over time in conjunction with wear. The use of 
cutting fluid yielded significant gains in part roughness, reaching Ra 
values lower than under dry testing, even after tool wear. In some tests 
(e.g., at 100 and 200 m/min cutting speeds), the roughness in both al-
loys remained practically constant until the end of the tool life. Under 
these conditions, the roughness remained practically constant during 
the tests and at values close to 0.30 μm. 
At a 200 m/min refrigerated cutting speed in material with Nb 
addition, a more atypical roughness behaviour could be observed in 
comparison with the other tests. This behaviour may have been influ-
enced by micro-chipping in the cutting edge or by abnormal wear of this 
cutting edge, resulting in a Ra close to 0.60 μm. At a 300 m/min cutting 
speed, roughness appeared to be slightly higher than in the other tests. 
This is understandable because at these speeds, the fluid has difficulty 
exercising its lubri-cooling function with the same efficiency as at other 
cutting speeds This is because the dynamics of the process derived from 
Fig. 17. Mean arithmetic deviation (Ra) as a function of removed material volume. 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
12
the high cutting speed lead to higher roughness. The cutting fluid 
radically changed the wear mechanisms of the tool during HCWCI 
machining, as described later. In general, the edges under dry cutting 
were more greatly affected by the abrasion phenomenon than when 
cutting fluid was employed. Abrasion creates grooves in the tool edge, 
particularly when machining materials that have a high volumetric 
fraction of carbides, such as HCWCI; however, these deeper grooves 
hardly appeared when cutting fluid was used. The geometry of the tool 
edge may have been transferred to the surface of the part at the time of 
machining. Therefore, the grooves generated by abrasion during dry 
cutting likely increased the intensity of the part roughness, but this did 
not occur in tests with cutting fluid. Even after the tool suffered wear, it 
still had a smooth appearance with no grooves, meaning roughness was 
reduced as compared to under dry testing. 
Profile total height (Rt). 
Fig. 18 shows the total height of the profile (Rt) obtained in all tests 
at the end of its useful life. In general, the roughness evolution Rt was 
very similar to the parameter Ra, increasing as the tool wear increased. 
It could also be observed that the cutting fluid significantly reduced the 
Rt values on the machined surfaces. 
For tests under dry cutting conditions, the roughness showed Rt 
values of between 5 and 8.5 μm. Rt roughness in HCWCI machining was 
hardly affected by the cutting speed. Instead, the primary factor was the 
wear derived from the cutting tool. Although in most tests, the Nb added 
material reached lower roughness than that without Nb, this difference 
was small and, in some cases, non-existent. Unlike the Ra parameter, 
carbide breakage (previously mentioned as a possible consequence of 
lower Ra values in Nb added materials) may have a lower impact on Rt 
roughness. This is because, with the total height of the profile measured, 
carbide breakage may not have as direct an effect as does Ra. It may 
interfere with the arithmetic mean butit will not accentuate the valleys 
and peaks generated during machining. Fig. 19 shows the evolution of Rt 
roughness as a function of the volume of removed material. 
Roughness remained practically stable under lubri-cooling condi-
tions. This indicated that, despite the wear of the cutting tool, the cutting 
edge did not generate grooves with sufficient intensity such that they 
transferred to the part, thereby increasing Rt roughness. Because it 
represents the total height of the profile, the Rt parameter may also have 
been mitigated by tool wear, thus de-characterising the tool tip radius 
and increasing the contact surface between the tool and part. Eventual 
spikes that developed on the machined surface may have been elimi-
nated by other parts of the tool, considering that the contact between the 
tool and part increased as the tool wear increased. 
3.6.4. Analysis of wear mechanisms 
Interpretations of wear mechanisms are fundamental to determine 
how the cutting tool wears out when subjected to cutting forces. HCWCI 
imposes severe cutting conditions on cutting tools during machining 
primarily because of the high presence of Cr carbides anchored in a 
martensitic matrix, which form a rigid microstructure that hinders the 
cutting phenomenon. It is likely that the wear of the cutting tools is not 
only generated by a wear mechanism but by a combination of mecha-
nisms that together result in the wear of the cutting tool. Analyses were 
conducted at the ends of the tests. Therefore, other wear mechanisms 
may have acted on the cutting tool at intermediate moments at the be-
ginnings and ends of the tests, and that they cannot be identified at the 
end of the cutting tool’s life. Analyses of the wear mechanisms aimed to 
identify qualitatively the main wear mechanisms that acted at the end of 
the cutting tool’s life. 
Fig. 20 shows aspects of wear on cutting tools at a 100 m/min speed 
both under dry cutting conditions and with cutting fluid in HCWCI alloys 
with and without Nb addition. 
Under dry cutting, Fig. 20(a)–(b) shows that abrasion was the pre-
dominant wear mechanism. The abrasion marks were concentrated 
mainly on the flank of the cutting tool, but it was possible to observe 
more discrete grooves in the crater, which may have been generated by 
abrasion. 
A comparison of the same cutting parameters revealed that the ma-
terial with Nb addition exhibited more pronounced crater wear. It was 
determined that the greater crater wear in the Nb added material may 
have been associated with the longer exposure of the crater to a high- 
temperature environment. This is because, in general, the tool life in 
Fig. 18. Total profile height (Rt). 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
13
machining Nb added material was higher during dry cutting, thus 
increasing the cutting time and exposure of the tool. The exposure time 
and cutting speed were described by Boucha et al. [18] as the two main 
factors that influence the wear of pcBN tools when turning hardened 
materials. In our study, considering that the tools were exposed for a 
longer time to a high temperature in the machining of Nb added alloys, 
thus allowing diffusion wear in the tool/chip region, a greater crater 
wear in this situation was understandable. 
The tests with cutting fluid considerably reduced the abrasion marks 
on the tool flank, indicating that the cutting fluid may have reached the 
cutting zone, in turn reducing the number of grooves caused by alloy 
carbides on the tool flank. Some grooves can be seen in Fig. 20(c)–(d) in 
the cutting fluid tests, mainly in the region near the cutting edge and at 
the tool tip. These regions are located where the fluid had more diffi-
culty reaching the part/tool interface due to the chip formation and 
specific cutting pressure. 
At a 100 m/min cutting speed with the use of cutting fluid, in the 
material without Nb addition, the 30 min time was achieved with low 
flank wear. However, the Nb added material showed chipping at the 
edge with a cutting time of approximately 17 min, and this occurred 
even with low flank and crater wear. 
According to Sales et al. [23], the machining of hardened materials is 
a typical example in which the cutting fluid can be detrimental to the 
process. The cutting fluid should work only as a coolant for the tool, but 
the regular application process causes the fluid to reach all chip for-
mation zones, also cooling the workpiece. Therefore, the softening effect 
caused by the large amount of heat generated is not substantial, 
requiring more energy in chip formation. It also induces high cutting 
forces and generates high temperatures at the tool/chip interface. 
Because of the high hardness of the part material, the decrease in 
hardness caused by the heat generation process is fundamental to 
improve the process performance. The cutting fluid prevents this and 
can negatively affect the process. In our study, edge chipping may have 
occurred because of the high cutting force derived from the increase in 
the specific cutting pressure that was due to the decrease in process 
temperature when using cutting fluid. 
Fig. 21 shows images of the worn areas of a tool at a 200 m/min 
cutting speed under dry and fluid cutting conditions. 
Fig. 21(a)–(b) shows that under dry cutting conditions, the wear 
presented by the tools was similar to that found at a 100 m/min cutting 
speed. With the use of cutting fluid, the characteristic grooves of abra-
sion wear were less clear at the tool cutting edge, indicating that the 
cutting fluid contributed significantly to a reduction in abrasion wear. 
Discrete grooves were still noticeable at the extremity of the cutting edge 
where the cutting fluid had difficulty reaching. 
Despite the lower flank wear, when compared to testing under the 
same dry conditions, during lubri-cooling cutting, the crater wear 
increased considerably. This wear may have been exacerbated by a 
longer exposure time of the material, as the cutting fluid increased the 
life of the cutting tool. The heat generated at the tool/chip interface may 
have increased during the process with the use of cutting fluid, as 
mentioned by Sales et al. [23]. 
Fig. 22 shows images of the worn areas of the tool at a 300 m/min 
cutting speed under dry cutting conditions and with cutting fluid for 
both HCWCI alloys. 
At a 300 m/min cutting speed, with the high cutting speed applied to 
Fig. 19. Profile total height (Rt) as a function of the volume of removed material. 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
14
a material with high hardness (approximately 62 HRC) such as HCWCI, 
abrasion marks were noticeable during dry cutting. When cutting using 
cutting fluid, although more discreet, these marks could also be seen. 
Crater wear was intense in these cutting conditions, causing the tools 
to show some chipping in both dry cutting and lubri-cooling conditions. 
This chipping was primarily noticeable in the material with Nb addition, 
presumably because of the longer exposure time of these tools to high 
temperatures, with crater wear weakening the edge and causing 
Fig. 20. Aspects of wear at a 100 m/min cutting speed in HCWCI both with and without Nb addition under dry cutting conditions and with cutting fluid. 
Fig. 21. Wear aspects at a 200 m/min cutting speed on HCWCI with and without Nb addition both under dry cutting conditions and with cutting fluid. 
A.E. Silva et al.Wear 460-461 (2020) 203463
15
chipping. 
The use of cutting fluid in this condition contributed significantly to 
increasing the life of the cutting tool but had little effect on the wear 
mechanisms. Abrasion marks were noted in all regions of the cutting tool 
flank under the high cutting speed conditions, indicating that the effect 
of the cutting fluid at high cutting speeds may have been compromised 
because of its difficulty in reaching the cutting zone. 
In all dry cutting tests, a greater number of grooves were observed in 
the tools that machined the material with Nb addition. Poulachon et al. 
[24], in the machining of AISI D2 hardened steel with presence of M7C3 
carbides, indicated that the grooves in the edge of the cutting tool may 
be related to the size of the primary Cr carbides. This behaviour can be 
observed in Fig. 23. 
The distance between the grooves was smaller in the Nb added ma-
terial, as shown in Fig. 23(a). This distance may be related to the average 
size of the carbides, which is smaller in this material, as a function of the 
refinement of the alloy by the addition of Nb. The alloy without Nb 
addition, as shown in Fig. 23(b), presented large primary carbides in its 
microstructure. These carbides may have caused large grooves that 
joined as they formed, creating a smoother surface. Both the sizes of the 
grooves and average size of the carbides were not constant. However, a 
correlation may exist between them. 
4. Conclusion 
In this study, the Nb carbides were formed in compact morphology 
Fig. 22. Wear aspects at a 300 m/min cutting speed on HCWCI with and without Nb addition both under dry cutting conditions and with cutting fluid. 
Fig. 23. Distance between grooves in relation to carbide size. 
A.E. Silva et al. 
Wear 460-461 (2020) 203463
16
and homogeneous dispersed into the alloy. In addition, most Nb carbides 
appeared in an integrated form with Cr carbides. 
The Nb addition promoted the refinement of the alloy, reducing the 
average carbide size and CVF from 27.3% to 24.83%. However, the alloy 
remained hypereutectic even after the Nb was added. 
The material hardness remained homogeneous throughout the 
sample after the heat treatment of tempering and quenching, indicating 
that the material had high hardenability. 
The Nb addition improved the machinability of HCWCI primarily 
during dry cutting. It also increased tool life and achieved significant 
gains. 
Under a 200 m/min cutting speed with the use of cutting fluid, the 
best results in terms of volume of the removed material were achieved, 
with the resulting roughness being similar to that obtained at a 100 m/ 
min cutting speed, thus indicating a more efficient machining condition. 
Abrasion was the major wear mechanism found in the cutting tools 
when machining HCWCI in both alloys. This mechanism was generated 
mainly by the M7C3 type primary Cr carbides. The grooves developed in 
the tools matched the sizes of the primary carbides presented in the 
microstructure of the material part. 
The cutting fluid positively contributed to the tool life, achieving 
gains in tool lifetime of approximately 200% in some cases. In addition, 
the cutting fluid generated reductions in roughness by as much as 75% at 
the ends of the tests, generating low roughness values even at the end of 
the cutting tool life. 
CRediT authorship contribution statement 
Anderson Edson da Silva: Conceptualization, Methodology, Vali-
dation, Formal analysis, Investigation, Resources, Data curation, 
Writing - original draft, Visualization, Project administration, Funding 
acquisition. Ismael Nogueira Rabelo de Melo: Validation, Writing - 
original draft, Visualization. Ivete Peixoto Pinheiro: Validation, 
Writing - original draft, Visualization. Leonardo Roberto da Silva: 
Conceptualization, Methodology, Validation, Formal analysis, Investi-
gation, Resources, Writing - original draft, Visualization, Supervision, 
Project administration, Funding acquisition. 
Declaration of competing interest 
The authors declare that they have no known competing financial 
interests or personal relationships that could have appeared to influence 
the work reported in this paper. 
Acknowledgements 
This work was conducted with the support of the Coordination for 
the Improvement of Higher Education Personnel - Brazil (CAPES) - 
Financing Code 001, the Federal Centre of Technological Education of 
Minas Gerais (CEFET-MG) and SENAI Itaúna CETEF Marcelino Corradi. 
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	Characterisation and machinability of high chromium hardened white cast iron with and without the addition of niobium
	1 Introduction
	2 Experimental procedure
	3 Results and discussion
	3.1 Chemical analysis
	3.2 Hardness
	3.3 Microstructure
	3.4 CVF
	3.5 X-ray diffraction
	3.6 Analysis of HCWCI machinability
	3.6.1 Lifetimes of cutting tools
	3.6.2 Effect of Nb addition
	3.6.3 Roughness analysis
	3.6.4 Analysis of wear mechanisms
	4 Conclusion
	CRediT authorship contribution statement
	Declaration of competing interest
	Acknowledgements
	References