3.

748

Article

High-Power Diode Laser Surface

Transformation Hardening of
Ferrous Alloys

Artur Czupryński, Damian Janicki, Jacek Górka, Andrzej Grabowski, Bernard Wyględacz,
Krzysztof Matus and Wojciech Karski

Special Issue
Advanced Technologies of Welding, Surfacing, and Thermal Spraying of Modern Materials
Edited by
Dr. Artur Czupryński, Dr. Marcin Adamiak, Prof. Dr. Antonín Kříž and Dr. Tünde Anna Kovács

https://doi.org/10.3390/ma15051915
 materials
Article
High-Power Diode Laser Surface Transformation Hardening of
Ferrous Alloys
Artur Czupryński 1, * , Damian Janicki 1 , Jacek Górka 1 , Andrzej Grabowski 2 , Bernard Wygl˛edacz 1 ,
Krzysztof Matus 3 and Wojciech Karski 1

1 Department of Welding, Silesian University of Technology, Konarskiego 18A, 44-100 Gliwice, Poland;
damian.janicki@polsl.pl (D.J.); jacek.gorka@polsl.pl (J.G.); bernard.wygledacz@polsl.pl (B.W.);
wojciech.karski@polsl.pl (W.K.)
2 Institute of Physics—CSE, Silesian University of Technology, Krasińskiego 8, 40-019 Katowice, Poland;
andrzej.grabowski@polsl.pl
3 Materials Research Laboratory, Silesian University of Technology, Konarskiego 18A, 44-100 Gliwice, Poland;
krzysztof.matus@polsl.pl
* Correspondence: artur.czuprynski@polsl.pl

Abstract: A high-power direct diode laser (HPDDL) having a rectangular beam with a top-hat
intensity distribution was used to produce surface-hardened layers on a ferrous alloy. The thermal
conditions in the hardened zone were estimated by using numerical simulations and infrared (IR)
thermography and then referred to the thickness and microstructure of the hardened layers. The
microstructural characteristics of the hardened layers were investigated using optical, scanning
electron and transmission electron microscopy together with X-ray diffraction. It was found that
the major factor that controls the thickness of the hardened layer is laser power density, which
determines the optimal range of the traverse speed, and in consequence the temperature distribution
in the hardened zone. The increase in the cooling rate led to the suppression of the martensitic





 transformation and a decrease in the hardened layer hardness. The precipitation of the nanometric
Citation: Czupryński, A.; Janicki, D.; plate-like and spherical cementite was observed throughout the hardened layer.
Górka, J.; Grabowski, A.; Wygl˛edacz,
B.; Matus, K.; Karski, W. High-Power Keywords: laser surface transformation hardening; high-power diode laser; ferrous alloys; thermog-
Diode Laser Surface Transformation raphy; cooling rate
Hardening of Ferrous Alloys.
Materials 2022, 15, 1915. https://
doi.org/10.3390/ma15051915

Academic Editor: Antonio Riveiro

1. Introduction
Laser surface treatment methods have been intensively employed for the past three
Received: 8 February 2022
decades for the enhancement of surface properties of different cast iron grades [1–13].
Accepted: 26 February 2022
Much research has been devoted to improving the tribological properties of cast irons
Published: 4 March 2022
via a laser surface transformation hardening (LSTH) process [10–12,14–18]. In contrast to
Publisher’s Note: MDPI stays neutral conventional surface hardening processes such as induction and flame hardening, LSTH
with regard to jurisdictional claims in ensures excellent surface finish and very low distortion of the workpiece. In many cases,
published maps and institutional affil- this eliminates any need for post-treatment finishing [19]. The precise control of the heat
iations. input together with the ability to heat localized areas make LSTH a unique technology
for the fabrication of locally selected wear-resistant hard surfaces on complex-shaped
components [20]. In recent years, there has been significant interest in the application of
LSTH to increase the tribological performance of ductile cast irons (DCI) in both the as-cast
Copyright: © 2022 by the authors.
condition and after austempering heat treatment [14–17].
Licensee MDPI, Basel, Switzerland.
This article is an open access article
Many researchers have reported on the significant increase in the wear properties
distributed under the terms and
of DCI after the LSTH process [14,16]. The tribological properties of the laser hardened
conditions of the Creative Commons
layers on DCIs depend upon the layer hardness, which, in turn, is affected by the phase
Attribution (CC BY) license (https:// composition of the layer [16]. On the other hand, the microstructural characteristics of the
creativecommons.org/licenses/by/ laser hardened layer and also the layer thickness are directly determined by the thermal
4.0/). conditions in the hardened zone [17]. For successful laser hardening, it is necessary to form

Materials 2022, 15, 1915. https://doi.org/10.3390/ma15051915 https://www.mdpi.com/journal/materials

Materials 2022, 15, 1915 2 of 17

a relatively homogeneous austenite structure in the hardened zone during a very short
period of the laser-induced thermal cycle. This requires notably higher temperatures on
the hardened surfaces than those during conventional hardening to achieve a sufficiently
high diffusion rate for the homogeneous austenitic structure [18]. As a result, to obtain
the desired layer thickness and to avoid partial melting, precise control of the surface
temperature is needed. However, scant data are available on the effect of LSTH parameters
on the thermal conditions during the processing of DCIs and the resulting microstructure
of the hardened layers.
The main objective of this work was to assess the effect of the LSTH parameters on
the thermal conditions in the hardened zone and the resulting change in the thickness and
the microstructure of the surface-hardened layer. The LSTH process was performed on a
pearlitic DCI using a high-power direct diode laser having a rectangular beam with a top-
hat intensity distribution. The thermal conditions in the hardened zone were determined
by using infrared (IR) thermography and numerical simulations. The surface-hardened
layers were characterized in terms of their surface finish, microstructure and hardness.

2. Materials and Methods

The chemical composition of the DCI (EN-GJS-700-2) used in this study is listed
in Table 1. The DCI composition was measured using both a LECO GDS 500A optical
emission spectrometer (LECO Corporation, St. Joseph, MI, USA) and a LECO CS125 Carbon
Analyzer (LECO Corporation, St. Joseph, MI, USA). The microstructure of the as-received
DCI was composed of a pearlitic matrix (ferrite content <15%) and graphite nodules with
an average diameter of about 30 µm (Figure 1a). The average thickness of the cementite
lamellae was 160 nm (Figure 1b). The hardness of the DCI in the as-received condition was
290 ± 15 HV0.5. The DCI specimens in the shape of a rectangular plate (40 × 30 × 10 mm)
were ground to an average roughness (Ra ) of approximately0.05 µm. The surface roughness
measurements were conducted using a Surftest SJ-210 portable surface roughness tester
(Mitutoyo Corporation, Kawasaki, Kanagawa, Japan). Prior to obtaining the reflection
measurements and the laser hardening, the test plates were cleaned in acetone.

Table 1. Chemical composition of the investigated DCI substrate (wt%).

C Si Cu Mn Cr Ni S P Fe
3.52 2.62 0.80 0.24 0.02 0.04 0.008 0.016 balance

(a) (b)
Figure 1. (a) Optical image showing the microstructure of the as-received DCI grade EN-GJS-700-2;
(b) In-lens SEM image showing the perlite matrix.

The reflectivity of the DCI specimens at room temperature was measured using a
setup comprising an Ocean Optics PC2000-ISA-PC Plug in Fiber Optics Spectrometer
Materials 2022, 15, 1915 3 of 17

(wavelength range 500–1000 nm; resolution: ±0.5 nm), an ISP-REF Integrating Sphere and
a WS-1 diffuse reflectance standard (Ocean Optics Inc., Dunedin, FL, USA). Reflectance
measurements were made using both diffuse and specularly reflected light.
The thermographic measurements were carried out with the use of a FLIR A655SC
thermal camera (Teledyne FLIR LLC., Wilsonville, OR, USA). The camera works in the
7.5–14.0 µm wavelength range and a 16-bit dynamic range at a 50 Hz polling rate and has
an accuracy of ±2% of reading.
The hardening trials were carried out with a high-power direct diode laser (HPDDL)
with a rectangular beam with a top-hat intensity distribution in the slow-axis direction
(HPDDL, Rofin-Sinar Laser GmbH, Hamburg, Germany). The dominant wavelength of
HPDDL used was 808 nm. The HPDDL beam spot size was 1.5 × 6.6 mm (length × width).
The focal plane of the laser optics was located at the DCI specimen and the slow HPDDL
beam axis was set to be perpendicular to the scanning direction. Argon (at a flow rate
of 10 L/min) was used during all hardening trials to prevent oxidation of the surface.
Additional details of the HPDD laser source have been previously given [21]. The selected
hardening conditions are listed in Table 2. The selection of the optimal range of processing
conditions was aimed at avoiding the finish grinding operation in potential industrial
applications. For this reason, the occurrence of partial surface melting was not accepted.

Table 2. Selected processing conditions of laser surface transformation hardening.

Processing Laser Power Traverse Surface Layer

Laser Power Heat Input 1 Remarks/
Condition Density Speed Hardness Thickness 2
(W) (J/mm) Quality 3
No./SHB No. (W/mm2 ) (m/min) HV0.5 (µm)
H1 0.1 240 867 ± 60 458 ± 14 PM
H2 400 40.4 0.2 120 755 ± 32 61 ± 11 NM
H3 0.3 80 290 ± 15 - NH
H4 0.2 180 841 ± 54 437 ± 15 PM
H5 600 60.6 0.4 90 724 ± 29 192 ± 11 NM
H6 0.6 60 689 ± 31 78 ± 10 NM
H7 0.4 120 796 ± 58 393 ± 16 PM
H8 800 80.8 0.6 80 690 ± 28 196 ± 12 NM
H9 0.8 60 684 ± 30 92 ± 11 NM
1 2 3
defined by the ratio of the laser power and the traverse speed; measured in the center of SHB; assessed in
terms of the surface condition: PM—partial melting, NM—no surface melting, NH—no hardening.

Single-pass hardened beads (SHBs) were examined by optical microscopy (Eclipse

MA100, Nikon Instech Co., Ltd., Tokyo, Japan), scanning electron microscopy (SEM, ZEISS
SUPRA 35, Carl Zeiss Microscopy GmbH, Jena, Germany), transmission electron mi-
croscopy (FEI Titan 80/300, Scientific and Technical Instruments, Hillsboro, OR, USA) and
X-ray diffraction (XRD). SEM investigations were conducted using both secondary electrons
and in-lens imaging modes. For both the SEM and optical investigations, the specimens
were etched with 4% Nital. The thickness of the hardened layer was measured using Nikon
image processing software on optical and SEM images taken from the SHB cross-sections.
The TEM investigations were performed using the conventional TEM and scanning trans-
mission electron microscopy (STEM) modes. STEM observations were carried out using a
high-angle annular dark-field (HAADF) detector. TEM samples were prepared using an
FIB technique using an SEM/FIB DualBeam microscope (FEI Helios NanoLab 600i FIB,
Scientific and Technical Instruments, Hillsboro, OR, USA).
The phase composition of the hardened layers was examined by using a PANalytical
X’Pert PRO MPD X-ray diffractometer (Malvern Panalytical, Almelo, The Netherlands).
XRD patterns were obtained using Co-Kα radiation (λ = 0.179 nm).
Materials 2022, 15, 1915 X’Pert 4 of 17
α λ

The hardness measurements were performed with a Wilson Wolpert 401 MVD micro-
hardness intender (Wilson Wolpert Instruments, Aachen, Germany). The measurements on
the hardened surface and the polished SHB cross-sections were conducted using 500 g and
100 g loads, respectively.
The laser hardening process was simulated by means of a nonlinear 3D transient
thermal simulation in Visual Weld 16 (Sysweld core-ESI Group). The material database con-
sisted of specific heat and thermal conductivity over the range of 20 ◦ C to the melting point.
– – –
The simulations take into consideration solid-state transformation with the Johnson–Mehl–
–
Avrami–Kolmogorov equation for diffusive transformations and the Koistinen–Marburger
equation for non-diffusive transformation. The equation parameters were chosen to fit the
simulated TTT diagram to the experimental TTT diagram. A 3D Model with 691,620 el-
ements and 187,640 nodes was developed (Figure 2). The dimensions of the developed
3D model were consistent with those of the specimens used in the experimental stage.
The model was prepared in such a manner that where the simulated laser processing took
place, hexa elements with a 0.05 × 0.05 × 0.05 mm size were used, and the rest of the
model was composed of tetra elements with element size increasing with the distance
from the processed area. Only half of the planar symmetrical sample was simulated to de-
crease computational complexity. This did not affect the simulation accuracy. The thermal
boundary conditions were set to free air with a temperature of 20 ◦ C exchange on the outer
surfaces of the model excluding the symmetry plane. The process was simulated with
the usage of a modified load 3D double ellipsoidal heat source proposed by Kik [22]. The
double ellipsoidal heat source dimensions were calibrated to achieve the 800 ◦ C isotherm
dimensions equal to the dimensions of the hardening zone. The heat source parameters
were as follows: length of 1.6 mm, depth of 0.1 mm and load zone width of 6.6 mm. The
results of reflection measurements (Figure 3) were taken into account during the selection
–
of energy input. The final heat source efficiency was in the 0.5–0.6 range.

Figure 2. The developed 3D model and meshing arrangement.

Materials 2022, 15, 1915 5 of 17

Figure 3. Spectral reflectance curve for a test plate of the DCI having an average roughness (Ra ) of
approximately 0.05 µm.

3. Results
3.1. Surface Condition Analysis
Figures 4 and 5 show the appearance of the laser hardened surface of the selected
DCI in different processing conditions. The results of the hardness measurements on the
hardened surfaces are listed in Table 2. The laser power level, determining the laser power
density, has a direct impact on the optimal range of the traverse speed. With an increasing
laser power density, the lower limit of the traverse speed, which ensures avoidance of
partial melting, is notably increased (Table 2). The partial melting occurred in the regions
directly adjacent to the graphite nodules and led to the formation of ledeburite eutectic
structures on the processed surface (Figures 4a and 5a). The presence of ledeburite eutectic
regions increased the overall surface hardness of the processed layers (Table 2). Note
that, to avoid finish grinding operations, surface melting is undesirable in transformation
hardening. The surface hardness of the SHB fabricated in the optimal range of parameters
(providing no surface melting) was in the range of 755 to 684 HV0.5. It should be noted that
with increasing traverse speed, the maximum surface hardness decreased. To understand
the effect of the processing conditions on the microstructure and the hardness of the laser
surface-hardened layers on DCI, the next sections present detailed investigations on the
microstructure and hardness of the SHBs produced under processing condition no. H2 and
H8, as well as the thermal conditions in the hardened zone under those two processing
conditions (Table 2).

(a) (b)
)
Figure 4. SEM images of the hardened surface at laser power of 400 W and traverse speed of:
(a) 0.1 m/min; (b) 0.2 m/min.
Materials 2022, 15, 1915 6 of 17

(a) (b)

Figure 5. SEM images of the hardened surface at a laser power of 800 W and a traverse speed of:
(a) 0.4 m/min; (b) 0.6 m/min.

3.2. Macro and Microstructural Analysis

Figure 6a,b show the cross-sectional macrographs of the SHBs produced under process-
ing conditions no. H2 and H8, respectively (Table 2). The corresponding low magnification
optical micrographs of the central area of the SHB cross-sections are presented in Figure 7.
Optical microscope images taken from the undersurface area of the above-mentioned SHBs
are shown in Figure 8. XRD patterns for the investigated DCI in the as-received condition
and after the LSTH under processing conditions no. H2 and H8 are presented in Figure 9a–c,
respectively. ––

(a)

(b)

Figure 6. Optical macrographs of the SHBs no. (a) H2; (b) H8 (Table 2).
Materials 2022, 15, 1915 7 of 17

(a) (b)

Figure 7. Low magnification optical micrographs of the SHB no. (a) H2; (b) H8 (Table 2). Dashed
lines indicate an approximate boundary between the hardened layer and the core material.

(a) (b)

Figure 8. Optical micrographs showing the undersurface area of the SHB (in the center of the bead)
no. (a) H2 (hardened layer and transition boundary) and (b) H8 (hardened zone) (Table 2).

The micrographs in Figure 7 suggest that the thickness of the hardened layer was
approximately 60 and 200 µm for conditions no. H2 and H8, respectively. Based on the
optical micrographs presented in Figure 8, one can observe that the microstructure of the
hardened layer is composed of graphite nodules and the martensitic matrix. The XRD
analysis indicated the additional presence of the retained austenite phase in the matrix of
the hardened layer produced under processing condition no. H8 (Figure 9c). Note that
the retained austenite fraction in the hardened layer produced under processing condition
no. H2 was negligible (Figure 9b). Moreover, XRD patterns from all hardened layers gave
peaks that can be identified as belonging to the cementite phase.
Materials 2022, 15, 1915 8 of 17

(a)

(b)

(c)
Figure 9. XRD patterns of the (a) as-received DCI and after the LSTH under conditions no. (b) H2 and
(c) H8 (Table 2).

SEM investigations (Figures 10 and 11) indicated that the hardened layer can be di-
vided into two sub-regions: the region of complete dissolution of cementite lamellae and the
region of partial dissolution of cementite lamellae. Directly beneath the hardened surface,
there was a region exhibiting a complete dissolution of cementite lamellae (Figure 11a). The
thickness of this region was approx. 22 and 135 µm in the SHB no. H2 and H8, respectively.
The microstructure of this region was composed of the graphite nodules and the martensitic
matrix (the martensitic/austenitic matrix in the case of SHB no. H8). The region of the
hardened layer adjacent to the core material had a matrix that was, in general, composed
of partially dissolved cementite lamellae and fine martensite laths (Figure 11b,c). The
thickness of the region of partial dissolution of cementite lamellae was approx. 39 and
61 µm in the SHB no. H2 and H8, respectively.
Materials 2022, 15, 1915 9 of 17

(a) (b)

Figure 10. In-lens SEM images taken from the undersurface area of SHB no. (a) H2 and (b) H8
(Table 2).

(a) (b)

(c)

Figure 11. In-lens SEM images taken from the SHB no. H8 (Table 2) showing: (a) the region of
complete dissolution of cementite lamellae; (b) the region of partial dissolution of the cementite
lamellae; (c) a boundary between the region of partial dissolution of the cementite lamellae and the
core material.

TEM investigations of the region of complete dissolution of the cementite lamellae

in the hardened layer (Figures 12 and 13) revealed the presence of nanometric plate-like
and spherical precipitates of cementite. The thickness of the cementite plates was close to
Materials 2022, 15, 1915 10 of 17

40 nm (Figure 12c). Both types of nanometric cementite precipitates were observed at the
martensite lath boundaries. Further research will be required to clarify the mechanism of
cementite precipitation during the investigated LSTH process.

(a) (b)

(c) (d)

Figure 12. (a) Dark-field and (b) bright-field STEM HAADF images showing a distribution of plate-
like (P) and spherical (S) cementite precipitates formed in the matrix of the SHB no. H8 (the region of
complete dissolution of the cementite lamellae); (c) Bright-field TEM image taken from a marked area
in (a); (d) SAED pattern obtained from region (1) in (a).
Materials 2022, 15, 1915 11 of 17

(a) (b)

(c) (d)
Figure 13. (a) Dark-field and (b) bright-field STEM HAADF images showing spherical (S) cementite
precipitates formed in the matrix of the SHB no. H8 (the region of complete dissolution of the
cementite lamellae); (c,d) SAED patterns obtained from regions (1) and (2), respectively, in (a).

3.3. Hardness Analysis

The hardness depth profiles of the SHBs no. H2 and H8 are reported in Figure 14. In
general, the hardness profiles correspond well to the microstructural changes discussed
in Section 3.2 and confirmed the previously mentioned thicknesses of the hardened layer.
Moreover, the hardness test data confirmed the decrease in the surface layer hardness
with increasing traverse speed mentioned in Section 3.1. Note that the subregions of the
hardened layer, i.e., the regions of complete and partial dissolution of the cementite lamellae,
have comparable hardness. This leads to a uniform hardness distribution throughout the
hardened layer. The average hardness values for the SHBs no. H2 and H8 were 755 ± 5 and
651 ± 15 HV0.1, respectively.
Materials 2022, 15, 1915 12 of 17

Figure 14. Hardness profiles of the SHBs no. H2 and H8 (Table 2).

3.4. Examination of Thermal Conditions in the Hardened Zone

The temperature distributions on the surface of the hardened specimens recorded by
the IR thermography and calculated for processing conditions no. H2 and H8 (Table 2)
are compared in Figure 15. The corresponding temperature profiles across the hardened
zone are shown in Figures 16 and 17, respectively. Figures 16 and 17 show that the
calculated temperature profiles match well with that recorded by the IR thermography,
indicating the reliability of the numerical model used. Figure 18a,b present the calculated
temperature fields in the cross-sections of the SHB no. H2 and H8, respectively. The
calculated temperature distributions beneath the surface of the SHB under the above
processing conditions are presented in Figure 19.

Processing condition no.: H2 Processing condition no.: H8

Figure 15. Temperature fields on the top surface of the laser surface-hardened specimens obtained via
IR thermography (upper part) and thermal simulation (lower part). Processing conditions: H2 and
H8 (Table 2). Selected emissivity value, E = 0.17.
Materials 2022, 15, 1915 13 of 17

(a) (b)

Figure 16. Temperature profiles for the hardened surface in (a) the longitudinal (along the SHB cen-
terline) and (b) traverse directions (along the slow-axis of the HPDDL beam). Processing conditions
no. H2 (Table 2).

(a) (b)

Figure 17. Temperature profiles for the hardened surface in (a) the longitudinal (along the SHB cen-
terline) and (b) traverse directions (along the slow-axis of the HPDDL beam). Processing conditions
no. H8 (Table 2).

Figure 18. Calculated temperature fields in the cross-section of the SHB (a half of the bead) no.:
(a) H2, (b) H8 (Table 2).
Materials 2022, 15, 1915 14 of 17

Figure 19. Calculated temperature distributions beneath the surface of the SHB (along the central
vertical plane) under processing conditions no. H2 and H8 (Table 2).

The thermal data indicated that the maximum temperature on the hardened surface
(on the SHB centerline) was about 1000 and 1100 ◦ C for processing conditions no. H2 and
H8, respectively. The average heating rate for the above two cases was 2300 and 8500 ◦ C/s,
respectively. The time spent above the critical temperature (i.e., the matrix structure
becomes entirely austenitic; assuming approximately 800 ◦ C) for processing conditions
no. H2 and H8 was 0.33 and 0.14 s, respectively. The average cooling rate within the
temperature range –of 800–500 ◦ C for processing conditions no. H2 and H8 was 1300 and
3400 ◦ C/s, respectively. The thermal data errors due to simplification of the modelled
process, limited mesh density, idealisation of the simulation and variety of emissivity
of the sample can be present. Although small deviations between the simulation and
thermographic experimental measurements are present, their convergence is high, which
increases the confidence of investigated thermal parameters.

4. Discussion
The laser power density, determining the optimal range of the traverse speed (ensuring
no surface melting), had a direct impact on the hardened layer thickness. The power density
for processing conditions no. H2 and H8 was 40.4 and 80.8 W/mm2 , respectively. Despite
a notably lower HI level (80 J/mm), the processing condition no. H8 provided an almost
3.5 times higher thickness of the hardened layers in comparison to that achieved under the
processing condition no. H2 (HI = 120 J/mm). Similar relationships between the processing
parameters and the hardening depth have been reported for LSTH of different grades of
steel [23,24]. It should be noted that the maximum hardness produced in the hardened layer
was affected by the traverse speed. The increase in the traverse speed led to a reduction in
the maximum layer hardness. Comparing the surface hardness of the SHBs no. H2 and
H8 shows that the lower traverse speed (0.2 and 0.6 m/min for H2 and H8, respectively)
provided about 10% greater surface hardness (Table 2). This trend of decreasing hardness
with increasing traverse speed is associated with an increase in the cooling rate and the
resulting suppression of the martensitic transformation. The lower hardness of SHB no. H8
(651 HV0.1, Figure 14) in comparison to that of SHB no. H2 (755 HV0.1) is attributed to the
presence of the retained austenite phase. The suppression of the martensitic transformation
under non-equilibrium cooling conditions has been reported in several works on the casting
and surface treatment of different cast iron grades [25,26]. It indicates that, from the point of
view of the maximum surface hardness, the lower traverse speeds, ensuring lower values of
the cooling rate, are beneficial in producing more wear-resistant layers [27,28]. On the other
hand, to avoid surface melting with increasing power density, the elevation of traverse
speed is required. The above results show that both the laser power density (determined by
the laser power level and laser spot size used) and the traverse speed are equally important.
Materials 2022, 15, 1915 15 of 17

Hence, a proper combination of the above processing parameters is required to achieve

optimization of the hardened layer hardness and depth.
Considering the notably shorter thermal cycle induced under processing condition
no. H8 (the time spent above the critical temperature for processing condition no. H2 was
almost 2.5 times longer than that for H8), it can be concluded the main factor controlling
the hardening depth under the analysed two cases is the value of the temperature on
the hardened surface. This relationship is consistent with that previously reported for
LSTH of steel [23,29,30]. The explanation can be found in the fact that the carbon diffusion
coefficient is a strong function of the temperature. The structural changes in the hardened
layer are dependent on diffusion-controlled processes such as the transformation of perlite
to austenite and the homogenization of the carbon content in austenite. Based on the data
reported by Gegner [31], it can be assumed that there is a significant increase in the carbon
diffusivity in the face-centered cubic Fe-C austenite lattice with increasing temperatures in
the range of 1000–1100 ◦ C.
Comparing the microstructure of the SHBs (Section 3.2) with the calculated tem-
perature distributions beneath the surface of SHBs (Figures 18 and 19) suggests that the
minimum temperature required for the complete dissolution of cementite lamellae in the
hardened layer was about 970 ◦ C. In turn, the formation of the hardened region that con-
tained the partially dissolved cementite lamellae underwent treatment in the temperature
range of 880–970 ◦ C.

5. Conclusions
Surface-hardened layers were produced on ductile cast iron via laser surface transfor-
mation hardening. The laser source used was a high-power direct diode laser having a rect-
angular beam with a top-hat intensity distribution. The thermal conditions in the hardened
zone were examined by using numerical simulations and infrared (IR) thermography. The
results showed that the major factor controlling the thickness of the hardened layer is the
laser power density, which determines the optimal range of the traverse speed and in con-
sequence the maximum surface temperature. The hardening depth was found to increase
with increasing power density. To avoid surface melting with increasing power density,
the lower limit of traverse speed notably increases. The resulting increase in the cooling
rate led to a suppression of the martensitic transformation and a decrease in the hardened
layer hardness. In the investigated range of the laser power density (40.4–80.8 W/mm2 ),
the maximum hardening depth was approx. 200 µm. The hardness in the hardened layer
was increased by almost three times in comparison with the as-received condition.

Author Contributions: Conceptualization, A.C. and D.J.; methodology, A.C., D.J., J.G., A.G., B.W. and
K.M.; formal analysis, A.C., D.J., J.G., A.G. and B.W.; investigation, D.J., A.C., A.G., B.W., K.M. and
W.K.; resources, A.C. and D.J.; writing—original draft preparation, A.C. and D.J.; writing—review
and editing, A.C. and D.J.; visualization, A.C., D.J. and B.W.; supervision, A.C. and D.J.; project
administration, A.C. and D.J.; funding acquisition, D.J. and A.C. All authors have read and agreed to
the published version of the manuscript.
Funding: The research was funded by the Silesian University of Technology Rector’s habilitation
grant 10/050/RGH_20/1006 and pro-quality grant 10/050/RGJ_22/1029.
Institutional Review Board Statement: Not applicable.
Informed Consent Statement: Not applicable.
Data Availability Statement: Not applicable.
Conflicts of Interest: The authors declare no conflict of interest.
Materials 2022, 15, 1915 16 of 17

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