Surface microrelief and hardness of laser hardened and ultrasonically peened AISI D2 tool steel

Surface microrelief and hardness of laser hardened and ultrasonically peened AISI D2 tool steel

Surface & Coatings Technology 278 (2015) 108–120 Contents lists available at ScienceDirect Surface & Coatings Technology journal homepage: www.elsev...

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Surface & Coatings Technology 278 (2015) 108–120

Contents lists available at ScienceDirect

Surface & Coatings Technology journal homepage: www.elsevier.com/locate/surfcoat

Surface microrelief and hardness of laser hardened and ultrasonically peened AISI D2 tool steel D.A. Lesyk a, S. Martinez b, V.V. Dzhemelinskyy a, А. Lamikiz b, B.N. Mordyuk c,⁎, G.I. Prokopenko c a b c

Department of Laser Physics and Applied Technologies, National Technical University of Ukraine “KPI”, 37 Peremohy Ave., UA-03056 Kyiv, Ukraine Department of Mechanical Engineering, University of the Basque Country UPV/EHU, Alameda Urquijo s/n, SP-48013 Bilbao, Spain Department of Solids Acoustics, Kurdyumov Institute for Metal Physics, NAS of Ukraine, 36 Academician Vernadsky Blvd., UA-03142 Kyiv, Ukraine

a r t i c l e

i n f o

Article history: Received 17 April 2015 Revised 15 July 2015 Accepted in revised form 23 July 2015 Available online 27 July 2015 Keywords: Laser surface hardening Ultrasonic impact treatment Surface roughness Surface waviness Hardness Tool steel

a b s t r a c t This paper is focused on experimental analysis of the effects of laser surface hardening (LSH) combined with subsequent ultrasonic impact treatment (UIT) on the surface microrelief, hardness and microstructure in nearsurface layers of AISI D2 high-chromium, cold worked tool steel. The LSH provides fast heating of the nearsurface layers to the temperature above that of the phase transformation and temperature then rapidly cools by a self-quenching process. The formed heat affected zone is hardened thanks to the rapid heating/cooling process affecting microstructure, phase composition and carbide formation. Conversely, the UIT induces multiple impact loads providing severe plastic deformation of near-surface layers, and the hardening occurs by a dislocation mediated process. The optimal parameters of each process were determined to obtain maximum hardness and regular surface microrelief. Further, complete analysis of the effect of combined treatment (LSH + UIT) on the surface hardness, microhardness depth profile and surface microrelief was performed. Results show that the combined LSH + UIT process provides almost triple, double and a 10% increase in hardness in comparison with those of the initial, UIT-processed and LSH-treated states, respectively. The surface microrelief, waviness and roughness parameters were respectively diminished after LSH + UIT by approx. 50, 65, and 90%. XRD analysis was carried out after LSH and LSH + UIT processes, which showed essential α-Fe peak broadening due to the formation of microstrains (0.27% and 0.47%, respectively) and reduction in crystallite size (84 nm after LSH + UIT). Favorable compressive residual stresses (−205 MPa and −409 MPa, respectively) were also observed in the near-surface layers of ~350 and 80 μm thick, respectively. The obtained results demonstrate that the combined LSH + UIT process is a feasible surface treatment for the quality improvement of the tool steel components including both the surface microrelief and hardness characteristics. © 2015 Published by Elsevier B.V.

1. Introduction One of the most important problems solved by mechanical engineering is to enhance the operational properties of products and components, which work in extreme conditions at high motion speeds, cyclically variable temperatures, high specific pressures, and under the actions of abrasive and aggressive environments. The performance of this group of products is largely dependent on the quality of their surfaces. Frequently, traditional technological processes for surface finishing don't fully satisfy the ever changing requirements. Much attention is now given to investigate the high energy methods for surface finishing [1,2]. Currently, finishing treatments used either highly concentrated energy sources [3,4] or high energy multiple impacts inducing severe plastic deformation of the surface [5,6]. The laser surface processing ⁎ Corresponding author. E-mail address: [email protected] (B.N. Mordyuk).

http://dx.doi.org/10.1016/j.surfcoat.2015.07.049 0257-8972/© 2015 Published by Elsevier B.V.

[7,8] resulting in melting [9], transformation hardening [10] or severe plastic deformation [11] of metallic surfaces, the surface treatments using electron beam [12], and plasma arc [13], as well as surface mechanical attrition treatment [6], deep ball-burnishing [14], ultrasonic impact treatment (UIT) [15–17], and ultrasonic nanocrystal surface modification [18] are promising cost-effective methods for surface hardening, surface polishing and formation of regular microrelief. As compared to the traditional energy sources for thermal hardening laser surface hardening (LSH) used the focused laser beam has the advantage over the other techniques. Continuous action of high energy density laser beam irradiating a local area allows both heating the surface and nearsurface layers and rapid cooling by a self-quenching due to conduction from the substrate. Rapid heating and cooling of thin surface layers promote microstructural changes via phase transformations in many steel products [4,8]. Conversely, strain hardening at surface by severe plastic deformation (SPD) causes structural changes by increasing the dislocation density, grain refinement, and formation of residual macro-stresses [2,14]. However, the temperature increases during operational life may

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cause partial or complete degradation of this material properties induced by the surface hardening methods [19]. Thus, development of novel techniques is a very topical task. Combined thermo-straining methods may be a promising treatment to improve magnitude and stability of surface hardening/finishing effects. Thus, the static roller burnishing [14,20] or dynamic action by ultrasonic treatment [21] or shot peening [22,23] used after laser hardening treatment has shown good results. There are several publications that have reported on the characteristics of the structure and hardness of steel surfaces [24,25] after the laser treatment followed by room temperature ultrasonic treatment [26,27]. However, to the best of the authors knowledge data describing the effects of combined action of the laser surface hardening and subsequent SPD process is practically absent. The aim of this work is to determine the optimal regimes of the laser hardening process and ultrasonic impact treatment for surface finishing and hardening of AISI D2 tool steel. Particular attention was paid in establishing the links between the parameters of the combined laser surface hardening and ultrasonic impact treatment, on the one hand, and surface microrelief, and hardness coupled with microstructure in the treated near-surface layers, on the other hand.

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controller was used to maintain a constant temperature in the working area [29]. Laser surface hardening was carried out with a scanning velocity V = 1000 mm/s which produced the track 10 mm wide on the specimen surface (Fig. 1). The temperature range was between 900…1340 °C, the specimen feed rate was between 40…140 mm/min. Specimens were laser processed in the atmosphere. The main parameters trialed for the LSH regimes are shown in Table 2. When using continuous laser radiation two strategies can be adopted to heat the surface either using a strategy of constant power (Fig. 2a) or a strategy of constant temperature (Fig. 2b). Optimal temperatures for these strategies were determined by our previous studies [29,30]. In this paper, the latter strategy was used, and an automated control system was employed to change the applied power of the laser beam to maintain the processing temperature. Unlike the strategy of constant laser beam power (Fig. 2a), the strategy of constant temperature (Fig. 2b) allows avoiding undesirable surface melting in the sites where the surface relief of the specimen is markedly changed. To maintain the necessary temperature regime several parameters need to be calculated to correctly control the laser beam power. Thus, a permissible heat energy density of laser beam, Elb in [J/cm2], in the irradiation area of treated surface was calculated for continuous mode as [31]:

2. Experimental details

Elb ¼ CP=dlb S;

2.1. Material

ð1Þ

where, P is the power of the laser beam in [W], dlb is the diameter of the laser beam in [mm] and S is the specimen feed rate of the treated surface in [mm/min], and C = 6000 is a dimensionality coefficient. The interaction time of the laser beam is determined by the feed rate S of the laser beam on the treated surface. The obtained magnitudes of the energy density fall into the range between 20…90 kJ/cm2. Accounting for the absorption ability of the material, А = 0.8, and dimensions, hlb and llb, of rectangular-shaped laser spot produced by the laser beam the heating temperature for continuous energy supply was determined by absorbed power density of the laser beam [32–34]:

The specimens of AISI D2 tool steel (69 mm × 69 mm × 9 mm) were initially heated to 850 °С, then slowly (10 °C per hour) cooled in the furnace to 650 °C, and then removed and cooled in air. The chemical composition and mechanical properties of the tool steel are given in Table 1. The resulting structure was composed of the alloyed α-ferrite and carbide phases. 2.2. Treatment details 2.2.1. Laser surface hardening Before the use of combined treatment, which consisted of LSH and subsequent UIT treatment in this study, a wide range of the individual parameters of the processes were evaluated first. The preliminary temperature range of laser surface hardening without melting was determined in accordance with the ternary iron–carbon–chromium phase diagram [27,28]. A Kondia Aktinos B500 machine with numerically programmed control and total working area of 500 × 300 mm was used for LSH. Scanlab Hurry Scan 25 2D scanner with total scan area in focal plane of 120 × 120 mm was placed into the machine, which contains two rotatable mirrors that allow moving the laser beam across a rectangular area on the specimen surface. Laser radiation was transported by optical fiber to the scan head of a Rofin Sinar FL010 fiber laser operated in continuous mode with a maximum power output of 1 kW and a wavelength λ = 1.07 μm ± 10 nm (Fig. 1). The scanning velocity V of the laser beam was controlled using the software of the scanner [29]. The optical fiber ensured cylindrical spot of approx. 1 mm in diameter, dlb, and multimode energy distribution in the focal plane. An Impac Igar 12LO two-color pyrometer was used to measure the temperature in the zone irradiated with the laser beam, and a special proportional integral

W lb ¼

AP : hlb llb

ð2Þ

The calculated power density, Wlb, falls into the range of 103… 10 W/cm2 at interaction time between 0.42 and 1.5 s, which is known to provide the surface hardening without melting [35]. In this study, maximum depth of the heat affected zone (HAZ) was approx. 0.3…0. 4mm (Table 2). 4

2.2.2. Ultrasonic impact treatment Severe plastic straining of near-surface layers was induced by UIT carried out using the equipment described in detail elsewhere [36–38]. The equipment contained an ultrasonic generator power output of 0.3 kW and a frequency of 21.6 kHz, an acoustic system with a piezoceramic transducer and step-like horn, the impact head placed on the horn tip (Fig. 3). At the UIT process, the amplitudes, Аusv, of ultrasonic horn vibration used were 15 and 18 μm, the acoustic system was pressed to the treated surface with a static force of Fs = 50 N, and shifted along the specimen surface with the feed rate S = 600 mm/min (Table 3). The high frequency (fi ≈3 ± 0.5 kHz) impacts were produced

Table 1 Chemical composition and mechanical properties of AISI D2 tool steel. Mechanical propertiesa

Chemical composition, % C

Cr

Mo

V

Mn

Si

Ni

Cu

P, S

Fe

σY [МPа]

σU [МPа]

δ [%]

E [GPа]

μ

ρ [kg/m3]

HB

1.55

11.3

0.8

0.8

0.4

0.3

b0.4

b0.3

b0.03

~84

≥320

≥710

≥16

200

0.25

7700

210

a

σY is the tensile yield strength, σU is the tensile ultimate strength, δ is the relative elongation, E is the Young's modulus, μ is the Poisson's ratio, ρ is the density, and HB is the Brinell hardness number.

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Fig. 1. Block-scheme of equipment for laser surface hardening and scheme for scanning process with focused laser beam: V is scanning velocity, b is the track width, S is the specimen feed rate, α1/α2 is the scanning angle of mirror 1/mirror 2. View of the heat affected zone (HAZ) in the specimen cross-section and view of the resulting track.

by seven cylindrical pins of 5 mm in diameter, positioned in the head that was forcedly rotated during treatment with a rotation speed of n ≈ 8.8 rpm. Duration of the UIT process, τ, was varied from 60 to 240 s (Table 3). Mechanical energy applied to the treated surface by the ultrasonic impact treatment Eui can be expressed by the impact frequency, fi in [s]; the pin mass, mp in [kg], the kinetic energy supplied by the ultrasonic horn, Eus, and additional kinetic energy of the pin, Ea induced by the head rotation (Eq. (4)) [15,38]: Eui ¼

f i Еus f ¼ i ðEus þ Ea Þ: mp mp

ð4Þ

The main component of the kinetic energy of the impacting pin is induced by the vibrating ultrasonic horn, and is determined by the linear velocity of the pin, vp, which equals the average vibration velocity of the horn tip. The latter is directly connected to the frequency, fusv in [s-1], and amplitude, Ausv in [m], of the tip vibrations [16]: Eus ¼

mp vp 2 2 ≈2mp π2 f usv A2usv : 2

ð5Þ

The additional kinetic energy Ea depends on the rotation frequency, nh in [rpm], of the head, the rotation radius of the pins, Rp in [m]

(distance from the center pins to the head axis (Fig. 3)), the angular speed of the pin, ωp, and can be expressed by the formula [16]: 2 Ea ¼ α nh πRp mp :

ð6Þ

The calculated mechanical energies applied to the treated surface at the UIT process were 45 and 64 kJ for amplitudes Ausv of 15 and 18 μm. The temperature increase at the UIT process was experimentally measured using thermocouple embedded into the specimen subsurface. The obtained magnitudes did not much varied dependently on the UIT regime and did not exceed approximately 150 °C. 2.2.3. Methods for evaluation of surface microrelief, hardness and microstructure The surface microreliefs and HAZ cross-sections were investigated using Leica DCM3D microscope by means of confocal lens 10XLD. The surface hardness of the specimens was determined using a Computest SC tester at a load on indenter of 10 N, and depth profiles of microhardness in the near-surface layers were registered using a FM800 tester at a load on Vickers indenter of 0.5 N. The intensity of surface hardening was evaluated by the HRC values of surface hardness of the initial and treated material by the following in − Hin formula uhard = (Hhard μ μ )/H μ ⋅ 100 %.

Table 2 The processing parameters of the LSH regimes. No.

LSH1…6

LSH 7

LSH8

LSH9

LSH10

LSH11

LSH12

LSH13

LSH14

LSH15

Т [°С] S [mm/min] HAZ width [mm] HAZ depth [μm]

900/1050 40/90/140 – –

1200 40 9.5 170

1200 90 7.8 140

1200 140 7.6 80

1270 40 10.5 380

1270 90 10.2 310

1270 140 9.5 240

1340 40 10.1 290

1340 90 10.0 270

1340 140 9.5 230

T is the temperature heating, S is the specimen feed rate, HAZ is the heat affected zone.

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Table 3 The processing parameters of the UIT regimes. №

UIT1

UIT2

UIT3

UIT4

UIT5

UIT6

UIT7

UIT8

Аusv [μm] τ [s]

15 60

15 120

15 180

15 240

18 60

18 120

18 180

18 240

Ausv is the amplitude of horn vibrations, τ is the duration of the UIT process.

The analysis of the microstructure in the near-surface layers was carried out using Nikon Optiphot-100 optical microscope. The crosssections of the treated specimens were mechanically polished and chemically etched using a 4% Nital reagent to reveal the grain boundaries and carbides' distribution. The X-ray diffraction analysis of the surface layers of about 10…20 μm thick was performed on the specimens in initial state and after treatments of different types and regimes. The X-ray measurements were carried out using a DRON-3 M diffractometer using a Cukα radiation with a graphite monochromator. Changes in width of reflexes and their angular positions on the diffraction spectra after LSH and UIT were analyzed to evaluate the lattice microstrains and crystallite sizes (coherent scattering areas — CSAs) and the treatment induced macrostresses σR = σ1 + σ2. The true physical broadening β of reflections was determined by subtracting the instrumental width from the full width B measured using the centroid of the diffraction profile. To estimate the peak broadening the registered diffraction peaks were fitted with pseudo-Voigt function using OriginPro 8 software, and the width, intensity and the centroid position of the fitted peaks were obtained and analyzed. The measured broadening B is known to be composed of the contributions of the size D of CSAs and the mean lattice micro-strain η in the form of the Scherrer equation [39]: B cosθ ¼ ð0:9λ=DÞ þ η sinθ:

Fig. 2. Laser heat treatments used the strategies of constant power (a) and constant temperature (b).

ð9Þ

Thus, according to the Williamson–Hall method when B cos θ is plotted against sin θ, a straight line is obtained with the slope as η reflecting the lattice micro-strain and the intercept as 0.9λ = D corresponding to the crystallite size D. Before the application of the surface treatments the stress-free lattice spacing d0 was measured in the annealed specimen. The obtained magnitude of d0 was then used to estimate residual macrostresses

Fig. 3. Schemes of ultrasonic impact equipment (a) and loading unit (b), view of the treated track (c).

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on the base of the shift of diffraction peaks. The estimations of macrostress in the surface layers were carried out using the following expression [40]:

σ R ¼ ðσ 1 þ σ 2 Þ ¼ −

E ðd−d0 Þ ; μ d

ð10Þ

where (σ1 + σ2) — is the plane stress, which is parallel to the sample surface; d and d0 are the interplanar distances in the stress-free state and after treatment, μ is the Poisson's ratio. 3. Results The previous experiments showed that the optimal temperature range of 1200…1340 °C for heating the AISI D2 steel can be achieved at the specimen feed rates of 40…140 mm/min. View and dimensions of HAZ on the cross-sections of the processed specimens (Fig. 1, Table 2) show that increasing specimen feed rate results in the decrease in the absorbed energy of scanning laser beam and subsequently in the decrease of the HAZ dimensions. HAZ was almost invisible in the specimens treated with LSH at temperatures less than 1200 °C. At higher temperatures the HAZ became visible, and both the hardening depth and width increased at the temperatures increasing from 1200 °C to 1270 °C and then they slightly decreased at a temperature of 1340 °C (Fig. 4ab). In comparison with the initial state the laser hardened specimens demonstrate approx. 2.7 times increase in surface hardness (Fig. 4c) and approx. 2 times higher near-surface microhardness in the specimens cross-sections (Fig. 4d) almost independently on the surface temperature.

The results of this set of experiments (Fig. 4) allowed establishing the optimal parameters for laser surface hardening regime. The following parameters of the LSH process at laser beam scanning velocity of 1000 mm/s are believed to be the most effective (the LSH#11 regime in Table 2): the specimen feed rate of 90 mm/min; the width of the scanning track of 10 mm; and the self-quenching temperature Тhard = 1270 °С. The LSH process using the fiber laser promotes only insignificant changing in geometric parameters of the processed surface. Initial state of the surface microrelief (polished/unpolished) also plays an important role. Severe plastic straining induced by UIT promoted the formation of the regular microrelief on the treated surface. In comparison with the initial state the UIT process resulted in a decrease in the roughness parameter Ra (Fig. 5a) and an increase in the surface waviness Wa (Fig. 5b) and average 3D roughness Sa (Fig. 5c) parameters of the surface microrelief. The most noticeable reduction in the roughness parameter (Ra = 0.13…0.18 μm) was registered after the UIT processes for 60 and 120 s (Fig. 5a). The UIT process of a previously polished surface (Fig. 5a) allowed reducing the Ra by 3.5 times. But at the same time the increase in waviness Wa (in 3.5 times) and in average 3D roughness Sa parameters (in 5 times) was observed in comparison with the initial polished state. Conversely, the unpolished surface, which is characterized with much larger initial surface parameters, undergoes marked changes (Fig. 5a–c, back columns). All the geometric roughness parameters of the microrelief profile were significantly decreased after the UIT process, in particular the Ra, Wa, and Sa were reduced by 9.5, 1.8 and 2.3 times, respectively. The UIT process also leads to the hardening of the surface layer. The lowest values of the surface hardness were obtained after the UIT processing for 60 s, while the largest increment in Rockwell hardness (up to 29.5 HRC) was observed after UIT for 180 s at the amplitude of

Fig. 4. Changes of the hardening depth (a) and width (b), the surface hardness (c) and microhardness (d) depending on the specimen feed rate and surface temperature.

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Fig. 5. Effects of the UIT process on the change of the microrelief parameters; surface roughness Ra (a), surface waviness Wa (b), average 3D roughness Sa (c), and surface HRC hardness (d) depending on the amplitude of ultrasonic vibrations and treatment duration (* denotes unpolished initial surface).

the ultrasonic horn of Auv = 18 μm (Fig. 5d). In this case, the hardening intensity was of 50% increase when compared with the substrate. However, further increase in the UIT duration at both used amplitudes caused some diminution in the surface hardness due to overhardening that induced an exfoliation of the surface. Thus, the results of this set of experiments (Fig. 5) allowed establishing the optimal UIT regime in the sense of maximum hardness (1.5 times higher than that of the initial state) and the lowest roughness. The following parameters of the UIT process are believed to be most effective (the UIT#6 regime in Table 3): a load of ultrasonic vibration system of 50 N; rotation speed of the impact head of 8.8 rpm; the amplitude of ultrasonic vibrations of 18 μm; and treatment duration of 120 s, which provides a specific impact quantity on the treated track area of 400 imp/mm2. Once the optimal parameters of each of treatments (single LSH#11 and single UIT#6) were established, in this second phase of work, a complete analysis of the surface characteristics produced by the sequential LSH#11 + UIT#6 process was addressed. Fig. 6 shows the surface profiles, roughness and waviness of the microreliefs characterized by different effects of phase transformation induced by LSH, dynamic straining by UIT and combined action of the LSH + UIT process on the aforementioned geometric parameters. There is no significant change in the profile parameters at LSH. In particular, the LSH results only in small decrease in the roughness parameter Ra and the waviness parameter Wa (Fig. 6b), which can be explained with the formation of oxide film due to the reaction of the specimen surface heated by a laser beam with the atmosphere. A reduction in the roughness parameter Ra and the waviness parameter Wa of the treated surface (Fig. 6c) was achieved after the optimal treatment duration of 120 s, and higher amplitude of ultrasonic horn (of 18 μm) producing more intensive impacts, and simultaneous forced

rotation of the impact head producing the sliding impact of pins by the specimen surface — every impact has high transversal constituent of load. Combined laser-ultrasonic treatment (LSH + UIT process) resulted in the formation of the regular and only slightly rough microrelief thanks to a much higher surface hardness of the LSH-processed surface of specimen, which inhibits overstraining and formation of excessive grooves, peaks and valleys. Besides, aforementioned high transversal constituent of load promotes sliding character of multiple high frequency impacts that causes surface flattening at the UIT. Hence, the waviness and roughness parameters are improved (Fig. 6d) due to intensive surface plastic straining, fragmentation and removal of the oxide film formed during heating by laser beam. Extremely small microasperities height parameters Sa registered by 3D laser microscopy (Fig. 7d) also confirm general tendency — the combined LSH + UIT appears the most effective in the sense of the formation of regular surface microrelief. Fig. 8 shows approx. 50–90% percentage reduction in geometric parameters of surface microrelief after the combined LSH + UIT process. In comparison with the initial state this reduction in the roughness parameter was even higher than that of a single UIT process, viz. 8.5 and 6 times, respectively. The obtained significantly reduced waviness and microasperity height parameters of the surface microrelief after the combined LSH + UIT process provided favorable conditions to trap oil on the product surface, which in turn, would promote reduced friction behavior for both static and dynamic operational conditions, and can be useful in industrial applications. X-ray diffraction patterns obtained from the initial and two treated specimens are shown in Fig. 9a. The pattern of the initial specimen (pattern 0) contains lines of the α-iron, traces of the γ-iron, and lines of the Сr7C3 and (Cr,Fe)7C3 chromium carbides. The volume fraction of

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Fig. 6. Change (average values of parameters) of the initial unpolished surface profile (curve 1), roughness Ra (curve 2) and waviness Wa (curve 3) of untreated surface (a) and after LSH#11 (b), UIT#6 (c) and LSH#11 + UIT#6 (d).

austenite is less than 10%. The presence of carbide phases is inherent for steels of ledeburite class like the studied AISI 2D steel, and these chromium carbides of different sizes and spatial distribution in the iron matrix mainly provide different levels of hardness and wear resistance. Our experimental results are in concordance with the Fe–Cr–C ternary phase diagram [28], which indicates the formation of chromium and chromium/iron carbides at cooling from the processing temperature. The LSH (pattern 1) leads to the formation of a large number of iron and chromium oxides on the specimen surface, and these oxides manifest themselves with the appearance of additional diffraction peaks. Moreover, accounting for the fact that the intensities of these additional peaks

are higher than those of the iron peaks, it can be argued that the LSH induced oxide layer is of sufficiently large thickness. Single UIT process (pattern 2) does not lead to oxide formation; it just results in broadening and shift of diffraction peaks. The UIT process carried out after previous LSH (spectrum 3) results in a reduction of the oxide peaks intensities, in significant broadening of the iron peaks, which also shift towards smaller diffraction angles (Fig. 9b). A decrease in intensity of the oxides lines is associated with their destruction and removal from the surface at the UIT process. The broadening of the diffraction peaks is known to be associated with two factors: a decrease in the size of CSAs and growth of the lattice microstrains. Diminution of CSA size is

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Fig. 7. Topography of the surface microrelieves of unpolished specimen in initial state (a) and after LSH#11 (b), UIT#6 (c) and combined LSH#11 + UIT#6 (d) processes.

correlated with a decrease in grain/subgrain size. The lattice microstrain can be increased due to two factors: (i) the increased dislocation density, and (ii) dissolution of carbon in the iron lattice by the decomposition of carbides (mainly cementite). The estimated values of crystallite size (CSA) and iron lattice microstrains were obtained by analysis of the broadening of the registered X-ray peaks using Williamson–Hall method and Scherrer Eq. (9). Broadening of the iron peaks after the LSH process (see for instance Fig. 9b) was concluded to be primarily related with the formation of microstrains (~ 0.27%) in the crystal lattice thanks to rapid selfquenching and phase transformation. On the contrary, the broadening, which is the largest after the combined LSH + UIT process, is already composed with contributions of the crystallite refinement (CSA ~84 nm) and much higher value of the lattice microstrains (~0.47%). Fig. 9b compares angular positions of (310) diffraction peaks of α-iron for the diffraction pattern of initial and treated specimens. It is

Fig. 8. Reduction in geometric parameters of the surface microrelief of initial unpolished specimen after LSH#11, UIT#6 and combined LSH#11 + UIT#6 processes.

seen that after the LSH, UIT and combined LSH + UIT processes the peak is slightly shifted towards lower diffraction angles that indicate the formation of compressive stresses. However, the estimated magnitude of the residual stress is slightly lower for the case of the LSH processed specimen (−205 MPa) than for the other ones treated with single UIT process (− 376 MPa) and combined LSH + UIT process (−409 MPa). Generally, such effect is almost always observed in laser treated metallic materials, in which near-surface layers are rapidly heated to high temperatures and further abruptly cooled down inducing the phase transformation. The UIT process forms compressive stresses in the near-surface layers due to their severe plastic deformation. In the case of the combined LSH + UIT process, both effects are summed up resulting in higher stress magnitude of approx. –409 MPa, which reaches about 50% of the yield strength of the investigated steel (~ 840 MPa). Besides, the deformation heating, which occurs at UIT, may lead to temperature increase in near-surface layer up to 150–200 °C. These temperatures are usually used at low tempering of retained austenite to obtain maximum hardness values. Increasing temperature (and hardness) at the UIT process can impede further plastic straining, and hence the formation of the compressive stresses of higher values. It is very important to know the case depth and distribution of residual stresses which were formed at different surface treatments. A stepwise electro-chemical removal of the stressed surface layers allows registering the depth profiles of residual stress formed in the AISI D2 tool steel after different hardening processes (Fig. 10). Notice however that the relaxation induced by electro-chemical removal leading to the stress redistribution and additional strains' occurrence was not taken into account here. All the processes studied result in the formation of compressive stressed regions. The depth of the compressed regions formed by LSH is about 350 μm which is much larger than that induced by severe straining at the UIT process (approximately 100 μm). In both cases, at larger depths the compressive residual stresses are balanced by tensile residual stresses. The residual stress depth profile formed after the combined LSH + UIT process is more complex — it contains fluctuation around the depth about 50–60 μm.

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Fig. 9. Diffraction spectra (a) and diffraction maximums (310) (b) of AISI D2 tool steel in the initial state (0), after the LSH#11 (1), after the UIT#6 (2), and combined LSH#11 + UIT#6 (3) processes.

These critical depths observed in residual stress depth profiles (Fig. 10) are in good concordance with the heat affected zone (HAZ) and strain affected zone (SAZ) visible in the specimen cross-section labeled in microscopic images of microstructure (Fig. 12). Moreover, a good correlation is also observed between the microstructural features revealed in the subsurface layers of AISI D2 tool steel after different hardening treatments (Fig. 12) and microhardness depth profiles (Fig. 11b). It is of interest that the SAZ thickness depends on the initial state of treated material. In comparison with the SAZ formed in asannealed specimen (Fig. 12c) the UIT induced SAZ in the laser hardened specimen is naturally thinner (Fig. 12d) due to its higher hardness/ lower plasticity (Fig. 11b). In comparison with the initial state all the treatment methods used lead to the hardening of the investigated steel coupled with microstructural alterations in near-surface layers (Fig. 11a). A typical microhardness depth profile after LSH (1270 °C, 90 mm/min) is shown in Fig. 11b. The thickness of the LSH hardened layer (about 310 μm) (Fig. 11) practically coincides to that of HAZ (Figs. 1, 12b, Table 2), and a sharp difference between the microhardness magnitudes is observed for the surface layer (HAZ) and bulk material. As mentioned above, the heating and subsequent self-quenching at LSH respectively increase the HRC hardness and microhardness in 2.6 and 2.1 times in comparison with initial specimen (or bulk parent material). It is due to the formation of martensitic structure of fine grained iron with dispersed particles of chromium carbides. Surface plastic deformation by means of a single UIT process leads to an increase in the surface HRC hardness by 40… 50% and in microhardness Hμ by 40% that is seemingly originated thanks to the increasing dislocation density and grain/crystallite refinement. However, microstructural alterations at the UIT process are known to

Fig. 10. Residual stress depth profiles formed in AISI D2 tool steel after the LSH#11 (1), UIT#6 (2) and LSH#11 + UIT#6 (3) processes.

occur in rather thin surface layer of about 60 μm thick [15–18,36–38]. The highest values of the surface HRC hardness (about 58.5) and microhardness (7.8 GPa) were observed after the combined LSH + UIT process (LSH#11 at a working temperature of 1270 °C and the specimen feed rate of 90 mm/min, and subsequent UIT#6), i.e. the increase almost in 3 and 2.4 times is respectively registered with regard to the initial state (Table 4). Thus, the UIT carried out after the LSH process promotes

Fig. 11. The microhardness of AISI D2 tool steel measured at the depth of 40 μm (a) in initial state and after the UIT#6, LSH#10, LSH#11 and LSH#12, and after LSH + UIT#6 processes, and depth distributions of microhardness in the surface layers (b) after the LSH#11 (1), UIT#6 (2) and LSH#11 + UIT#6 (3) processes.

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Fig. 12. Light microscopy images of microstructure in subsurface layers of AISI D2 tool steel in initial state (a) and after LSH#11 (b), UIT#6 (c) and LSH#11 + UIT#6 (d–f) processes.

further 10…15% increase in the surface hardness and microhardness (Table 4, Fig. 11b) by additional strain induced refinement of martensitic grains and carbides in the surface layer. 4. Discussion The results of this work clearly indicate the beneficial effect of surface treatment, specifically by the combined LSH + UIT process, in enhancing both the surface relief parameters and microhardness of AISI D2 steel. The results further show that the main contribution to the roughness reduction is due to the UIT process whereas the increased microhardness can be microstructurally attributed to the formation of sandwich-like hardened layer with a favorable compressive residual stresses. The UIT induced work-hardened near-surface layer covers the transformation hardened layer caused by LSH. This combination of the reduced surface roughness and enhanced hardness allows us to conclude that the combined LSH + UIT process is a feasible surface treatment for the quality improvement of the tool steel components. Let's compare our observations regarding the surface roughness, residual stresses and hardness with the literature data. 4.1. Surface roughness Surface roughness is one of the crucial parameters affecting the resistance of metallic materials to fatigue, wear and corrosion. In comparison with the LSH, which has negligible influence the surface roughness, the UIT and LSH + UIT processes were observed to markedly reduce it (Figs. 6, 8, Table 4) similar to other SPD methods [2,14,41, 42]. It is not surprising, and it is in good agreement with conclusions

of a number of works [7,11,31,43,44] studied the effects of different laser treatments. These studies report that the effects of laser treatment on the surface roughness are strictly dependent on the laser intensity, and they can either lower or conversely enlarge the roughness. The largest changes occur after laser melting treatment (LMT) [7,31] and laser shock peening (LSP) treatment without protective layer which induces ablation of the treated surface [11]. Thus, the laser polishing was shown to cause the maximal roughness reductions when the energy density range was chosen to be sufficient to melt thin surface layer [31], and the reduced roughness of AISI 1045 steel specimens improved their fatigue behavior [44]. Laser gas assisted treatment was also shown to considerably lower the surface roughness of AISI H12 tool steel if the laser pulse intensity was low enough to avoid large scale surface ablation [43]. On the contrary to the LSH process, UIT appears more effective in the roughness reduction (Figs. 6, 8, Table 4) similar to the other SPD methods. Previous studies on shot peening (SP), deep rolling (DR) and LSP of titanium alloy [41] or stainless steel [42] have indicated that the only DR decreased the surface roughness. Similar results were found for the ball-burnishing which improved the surface quality of the AISI 1045 steel workpieces and favorably influenced their fatigue life and wear resistance [14]. The SPD methods were also shown to be effective even at combined treatments, and our data is in good correlation with this conclusion. Thus, the burnished surface of AISI 5140 steel was distinctly flattened in comparison with that after previous cryogenic turning [45]. The SP process changed the surface roughness of the laser melted low-carbon chromium steel significantly [23], and some hardening induced by LMT reduced the ability of SP to diminish the roughness of the laser hardened surface. At the same time, the increase

Table 4 Geometrical parameters and mechanical properties of the surface layer of AISI D2 steel. Treatment

Geometrical parameters Ra

Initial LSH#11 UIT#6 LSH#11 + UIT#6

Mechanical properties

Wa

Sa

Surface hardness

[μm]

[%]

[μm]

[%]

[μm]

[%]

HRC

[%]

2.60 2.56 0.42 0.30

– 1.5 83.8 88.5

1.66 1.56 0.8 0.54

– 6.0 51.8 67.5

3.91 3.83 2.46 1.98

– 2.0 37.0 49.3

19.6 52.1 28.4 58.5

– 156.8 44.8 200

Hμ0.05 [GPa]

hhard [μm]

σR [MPa]

3.2 6.75 4.5 7.8

– 310 60 310

0 −205 −376 −409

Productivitya [mm2/s]

– 15 7.5 15

Ra is the arithmetic mean deviation of the roughness profile (ISO4287), Wa is the arithmetic mean deviation of the waviness profile (ISO4287), Sa is the arithmetic mean height of the profile (ISO25178), HRC is the Rockwell hardness number, Hμ0.05 is the microhardness of the surface layer, hhard is the hardening depth, and σR is the residual macrostress in the surface layer. a Productivity is determined by the time needed to treat a given area.

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in the SP intensity resulted in rougher surface than those in initial and laser melted states. Our results on the LSH + UIT process also indicate that surface roughness can be essentially minimized at optimal regime observed in this study to be LSH311 + UIT#6 (Fig. 8, Table 4). 4.2. Residual stresses Residual stress is another important factor affecting the materials performance. For the process parameters used, the UIT process was found to induce the larger near-surface compressive residual stresses but the lower degree of work hardening as compared with the LSH process (Fig. 11b), although the depth of the affected near-surface layer was larger after the latter process (Fig. 12). These observations are in good agreement with the data regarding both the variety of laser treatments (LSH [1,7], LMT [1,7,8] and LSP [1,7,11]) and such SPD methods as SP [41], DR [4,19], and UIT [2,15]. All the known methods of surface SPD result in high compressive stresses in top surface layers. The main differences consist in the thickness of the stresses layer and stress gradient. In this sense, the compressive stresses formed at the UIT process are comparable with those after SP [2,41], ultrasonic shot peening [41], but DR and especially LSP produce much thicker compressed layers [19,41,42]. Considering laser treatments, we can conclude that the sign, magnitude and distribution of residual stresses in laser treated layers are very affected by the energy density and interaction time, which induce the appropriate temperature regime and one of the following core processes: the transformation hardening due to quick heating/cooling, the surface melting/quick solidification, and surface shock hardening. Tensile residual stresses can be formed after LSP of uncoated specimens due to severe surface melting [11], after the LMT of surface layers of nodular iron [46] or tool steel [7] as well as under slow transformation-free cooling of laser heated steel surface. Conversely, the LSH process of an AISI 4140 steel, which was similar to the LSH used in the present study, induced only surface heating without melting. It was reported to form both compressive residual stresses inside the outer surface layer, which consisted of martensitic phase, and compensative tensile residual stresses outside of the HAZ — the zone irradiated by the laser beam [47]. Those residual stresses were naturally induced by solid state transformations in HAZ, viz. the surface austenization and further quenching induced martensitic transformation. In our case, XRD patterns (Fig. 9a) contained only traces of γ-iron witness to negligible volume fraction of the retained austenite and predominance of martensitic phase after LSH and LSH + UIT, which cause compressive residual stresses (Fig. 10). The UIT process following after LSH even increases the stress magnitude in thin top surface layer (curve 3 in Fig. 10). The latter result is in good agreement with the recent research of combined effects of laser surface melting and shot peening on the fatigue resistance of 20CrMnTi steel [23]. It was indicated that a large and deep compressive residual stress field was formed by shot peening in the laser surface melted layer of the steel. 4.3. Microhardness In addition to smooth surface finish and beneficial compressive residual stresses, the microhardness coupled with microstructure is also a crucial parameter for a long-term performance of metallic materials. Laser surface hardening is known to be achieved either by melting a steel surface, by alloying it, by straining it with shock waves or by heating it without melting [7]. Surface SPD methods result in significant work-hardening due to high rate straining to high strain extents. The results obtained in this work indicate that the combined LSH + UIT process is most effective in surface hardening albeit that the separately used LSH or UIT processes also lead to essential hardening (Fig. 11). The sandwich-like surface hardened layer is formed by the combined LSH + UIT process (Fig. 11b). The transformation hardened layer induced by LSH is four times thicker (between 80–310 μm)

than the UIT induced work-hardened near-surface layer (of 40–80 μm thick) (Fig. 11b), and both these layers are distinctly visible in appropriate cross-section microstructures (Fig. 12). Hardness magnitudes (51.2 and 58.5 HRC) registered in this study for the AISI D2 steel specimen after LSH and combined LSH + UIT processes are higher than that recently reported for this steel treated with LSH (48 HRC) [48]. Similar transformation hardening of 1045 steel and W–Cr–V steel was also reported in [7]. Microhardness (7.8 GPa) obtained after combined LSH + UIT process is higher than that observed after LSH of AISI H13 steel (5.3 GPa) [49] or comparable with that of AISI D6 steel (~ 6– 8 GPa) [50]. A couple of works were dwelt with the laser melting treatment of chromium carbon steel [51] and tool steel [9]. The former steel was reported to have unchanged hardness after LMT, but the latter one was conversely hardened. The following two studies had reported on the combined treatments. Thus, intensive friction of laser melted chromium carbon steel at wear tests resulted in 1.5 times increase in hardness [52]. Additional hardening was observed after shot peening of the laser melted low carbon chromium steel [23] similar to the case of LSH + UIT (curve 3 in Fig. 11b). 4.4. Summarizing remarks A systematic study of the effects of single LSH and single UIT processes on geometric parameters of the surface microrelief, hardness and microstructure in near-surface layers of AISI D2 high-chromium tool steel was carried out. The optimal parameters were determined for each process to obtain maximum hardness and regular surface microrelief. The following parameters of the LSH process are concluded to be the most effective: laser beam scanning velocity of 1000 mm/s; the width of the scanning track of 10 mm; the workpiece feed rate of 90 mm/min; and the self-quenching temperature Тhard = 1270 °С. The most effective UIT process is concluded to have the following parameters: a load on the acoustic system of 50 N; rotation speed of the impact head of 8.8 rpm; the amplitude of ultrasonic vibrations of 18 μm; and treatment duration of 120 s, which provides a specific impact quantity on the treated track area of 400 imp/mm2. The solely used LSH#11 process resulted in sufficiently high hardening (52.1 HRC, HV = 6.75 GPa) of thick near-surface layer (~ 300… 350 μm) due to transformation hardening during thermal cycling. It also produced moderate compressive residual stresses (−205 MPa) in the heat affected zone of about 350 μm thick. However, its effect on the surface roughness was negligible. The UIT#6 process caused essential reduction in the surface roughness. It also provided lower hardening (28.4 HRC, HV = 4.5 GPa) of thin near-surface layer (~80…100 μm) due to work hardening during severe plastic deformation. However, the UIT produced twice higher compressive residual stresses (− 376 MPa) in the strain affected zone of about 100 μm thick. Once, the optimal regimes of single processes of the AISI D2 steel were established, a complete analysis of the effect of combined treatment (LSH+ subsequent UIT) on the surface hardness, the microhardness and residual stress depth profile and surface microrelief was performed. Results show that the surface microrelief in terms of height (Sa = 1.98 μm), waviness (Wa = 0.54 μm) and roughness (Ra = 0.3 μm) parameters was respectively diminished after LSH + UIT by approx. 50, 65, and 90%. The combined LSH#11 + UIT#6 process of the AISI D2 steel provides enhanced hardness (58.5 HRC, HV = 7.8 GPa), which is three times, twice and 10% higher in comparison with those of the initial, UIT-processed and LSH-treated states, respectively. This microhardness increase is microstructurally attributed to the formation of the UIT induced work-hardened near-surface layer which covers the transformation hardened layer caused by LSH. The XRD analysis was carried out after LSH, UIT and LSH + UIT processes, which showed essential α-Fe line broadening connected to the formation of microstrains (0.27%, 0.32% and 0.47%, respectively) and diminishment in crystallite size (84 nm after LSH + UIT). After the combined LSH + UIT process, the highest magnitudes of favorable compressive residual stresses

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(−409 MPa) and slightly lower one (−200 MPa) were observed in the near-surface layers. The case depths of the compressed regions formed by the UIT and combined LSH + UIT processes were about 80 and 350 μm, respectively. In all cases, at larger depths the compressive residual stresses are balanced by tensile ones. 5. Conclusions This paper investigates the effect of laser surface hardening and ultrasonic impact treatment on surface roughness and microhardness of AISI D2 high-chromium tool steel. The study reaches the following conclusions. (i) Analysis of the hardness magnitude and the depth and width of the laser hardened area allowed to choose the optimal LSH regime, which provides ~ 150% hardness increase due to transformation hardening but negligible reduction in the surface roughness. (ii) Determination of optimal UIT regime was based on the lowest surface roughness achieved. The optimal UIT resulted in severe plastic deformation leading to ~ 80% reduction in the surface roughness and ~50% hardness increase. (iii) A complete analysis of the surface characteristics and hardness produced by the sequentially applied LSH#11 + UIT#6 processes demonstrated that the combined laser-ultrasonic treatment is a feasible surface treatment for the quality improvement of the tool steel components including simultaneously beneficial the surface microrelief and hardness characteristics of sandwich-like hardened layer compressed by residual stresses. Further increase in the operational properties of the complexly treated AISI D2 tool steel parts can be expected. Particularly, the increase in the opportunities of economical use of lubricants to enhance the wear and friction behaviors can be appreciable.

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