Study of mechanical and tribological properties of Ti–6Al–4V alloy fabricated by powder bed fusion laser beam

Abstract

Powder bed fusion laser beam, as one of the most promising forming technologies, offers unmatched benefits over traditional processing, particularly in the production of Ti–6Al–4V. The influence of laser line energy density (LLED) on the forming surface, phase composition, micro-hardness, tensile characteristics and wear resistance of Ti–6Al–4V alloy were explored to disclose the evolution of mechanical and tribological properties. According to the findings, the LLED causes ‘depressions’ and ‘highlands’ between nearby scanning tracks. The phase composition did not alter appreciably as LLED increased. Micro-hardness and tensile characteristics increased initially, then declined, and the value of maximum micro-hardness and ultimate tensile strength were 388.17 HV_0.2 and 1197.5 MPa, respectively. Furthermore, when the LLED is 0.24 J mm^–1, the wear resistance is optimal under the aviation lubricant medium, with an average friction coefficient of 0.1505 and volume wear rate of 6.95*10⁻⁸ mm²*N⁻¹, and a wear mechanism of mild furrow wear and adhesion wear.

Keywords

Powder bed fusion laser beam (PBF-LB)Ti–6Al–4V laser line energy density (LLED)mechanical and tribological properties

1. Introduction

Ti–6Al–4V is widely available for the aerospace field, biomedical field, marine environment and petroleum metallurgy field due to its excellent characteristics of low density, distinguished specific strength, splendid corrosion resistance and high fatigue strength [1 –3]. Although titanium alloy has many advantages, it also has some drawbacks, which limit its further application and development [4 –6]. Because of the inherent characteristics of a hexagonal close-packed crystal structure with a less slip system, results in work hardening by traditional machining methods [7,8]. Besides, as a result of the titanium alloy's low elastic modulus and the workpiece's high clamping and stress deformation during machining, the workpiece's machining precision will suffer. Consequently, it is extremely difficult to manufacture Ti–6Al–4V alloy using the traditional machining method [9,10]. Compared with the traditional machining methods, additive manufacturing (AM) can manufacture near-net shape parts with privileged design freedom by using metals and their composites, which supply a new method to prepare Ti–6Al–4V alloy [11 –14]. As one of the most popular AM technology, powder bed fusion laser beam (PBF-LB) enables to fabricate high-performance alloy components in a layer-by-layer fashion through melting alloy powder with a high-energy laser beam [15 –17]. Attar et al. [18] contrasted the hardness and compression test results of titanium alloy prepared by PBF-LB with that fabricated by casting method and found that the PBF-LB titanium alloy has better performance in mechanical properties due to its fine grain size and martensitic phase composition. Bartolomeu et al. [19] compared the effects of hot pressing, casting and PBF-LB processes on the wear resistance and the hardness of Ti–6Al–4V alloy. The results showed that due to the fast cooling rate, wear rate and hardness value of the PBF-LB part were 21% lower and 14% higher compared to the casting part, respectively.

However, the mechanical and physicochemical properties of Ti–6Al–4V alloy need further improvement to meet some extreme working conditions. Khorasani et al. [20] investigated the influence of SLM processing parameters on mechanical properties of Ti–6Al–4V, and found that higher laser power and slower scan speeds resulted in higher hardness due to improved energy transfer and better melting pool quality. Chen et al. [21] studied the effects of energy input on the microstructural, electrochemical and wear-corrosion characterisation of Ti–6Al–4V alloy obtained by PBF-LB, and the results demonstrated that higher energy input can make a significant improvement in wear-corrosion. He et al. [22] Conducted the effects of build direction on mechanical property of samples with 0°, 45° and 90° three build directions manufactured by PBF-LB and found that 90° sample had the highest strength and elongation, compared with forged samples, the PBF-LB samples had longer creep ruptured time but lower ductility.

The effects of single-factor processing parameters (such as laser power, scanning speed, etc.) on the microstructure and mechanical properties of pure Ti–6Al–4V components fabricated by PBF-LB have been widely reported. Hence, this study comprehensively considers the synergistic effects of laser power and scanning speed, by designing a reasonable combination of processing parameters and then investigating the influence of different LLED on the mechanical and tribological characteristics of PBF-LB Ti–6Al–4V. All the noticeable observations have been analysed and interpreted appropriately.

2. Materials and methods

2.1. Specimens preparation

Gas atomised Titanium alloy (Ti–6Al–4V) powder with excellent sphericity and particle size distribution of 15∼53 μm was used in this work. Cubic specimens with a dimension of 11 × 11 × 8 mm and ‘dog-bone’ tensile plates with a gauge length of 28 mm, a width of 5 mm and the thickness of 2 mm were fabricated by the PBF-LB machine (XDM 120, Suzhou XDM 3D Printing Technology Co., Ltd., China) equipped with a single-mode fibre laser of 500 W and wavelength of 1064 nm, as shown in Figure 1. The laser beam is focused to a spot size of approximately 100 μm, the layer thickness is 30 μm and a bi-directional scanning strategy of 67° rotation between two contact layers was adopted to obtain low and uniform residual stress and strain. To avoid the oxidation of the Ti–6Al–4V samples during manufacturing, a continuous flow of high-purity argon was pumped to the processing chamber to achieve an oxidation-free environment. Based on a lot of previous experiments, the laser power (210∼270 W) and scanning speed (1000∼1600 mm s^–1) were considered as suitable process windows for Ti–6Al–4V fabricated by PBF-LB. The detail of PBF-LB parameters is listed in Table 1. Element composition of Ti-6Al-4V powder is listed in Table 2. To figure out the effect of the laser energy input on the manufacturing quality, the LLED (E) is defined by:

(1) where P is the laser power and V is the scanning speed.
Figure 1
(a) SEM micrograph of Ti–6Al–4V powders used in this paper; (b) size distribution of the powder feedstock; (c) photos of the cubic samples and tensile specimens and (d) dimensions of samples used in tensile test.

Table 1
PBF-LB processing parameters.

Ti–6Al–4V samples Laser power (W) Scanning speed (mm s^–1) Hatch distance (mm) Laser liner energy density (E: J mm^–1)

S1 240 1600 0.12 0.15

S2 240 1300 0.12 0.18

S3 210 1000 0.12 0.21

S4 240 1000 0.12 0.24

S5 270 1000 0.12 0.27

2.2. Specimen characterisation

Ti–6Al–4V samples	Laser power (W)	Scanning speed (mm s^–1)	Hatch distance (mm)	Laser liner energy density (E: J mm^–1)
S1	240	1600	0.12	0.15
S2	240	1300	0.12	0.18
S3	210	1000	0.12	0.21
S4	240	1000	0.12	0.24
S5	270	1000	0.12	0.27

Before microstructure observation, all samples were polished by SiC grinding paper with grit size up to 5000 and etched in Kroll solution composed of 96 ml H₂O, 4 ml HNO₃ and 1 ml HF. The scanning electron microscope (Gemini SEM 300, ZEISS, Germany) was used to examine metallographic specimens and investigate the fracture surfaces of the tensile samples for identifying the fracture mechanism of the PBF-LB Ti–6Al–4V samples. The surface topography of the PBF-LB samples was characterised by a Profilometer (MFD-D, RTEC Instruments, USA). The phase of different parameters of PBF-LB samples was identified by an X-ray diffractometer (D2 PHASER, Bruker, Germany) with Cu Kα radiation at the scan speed of 4° min^–1 from 10° to 90° (2θ). The roughness was measured by Surface Roughness Tester (SJ 410, Mitutoyo, Japan). The Archimedes drainage method was used to measure the relative density of samples according to the formula [10,23]:

(2) where is the relative density of the sample, is water density, is the weight of sample suspended and immersed in water, is the weight of samples in air.
2.3. Mechanical tests

The Vickers micro-hardness of the PBF-LB samples was measured by using a micro-hardness tester (SANS-CMT-5205, Wance, China) under a load of 200 g with the dwell time of 15 s. For each specimen, five measurements from random positions on the polished cross-section of the specimen were taken and the mean value was calculated. An electronic universal testing machine (WDW-100KN, Jinan Sida Testing Technology Co., Ltd, China) with an extensometer was used under room temperature for tensile tests with a tensile rate of 0.5 mm min^–1 according to the ASTM E8M. For each sample group, three tests were taken and averaged to obtain the mean value. Friction performance of the PBF-LB samples was measured by using a Multi-Function Tribometer (MFT 5000, RTEC Instruments, USA) with a load of 20 N and a frequency of 2 Hz for 30 min. The experiment was carried out at room temperature in the lubrication medium of aviation lubricating oil (BP2389, Eastman, USA). A Si₃N₄ ball with a diameter of 9 mm was selected as the counterpart material. The wear volume was calculated through the 3D profiler software (RTEC). The wear volume in this work was calculated by the formula [24,25]:

(3) (4) (5)

where is the wear rate; is the wear volume; is the sliding distance; is the applied load; is the radius of the Si₃N₄ ball; is the sliding speed; is the testing time; is the wear length and is the wear width.
3. Results and discussions

3.1. Surface morphology and porosity

Figure 2 exhibits the untreated surface topography of the PBF-LB samples along the printing direction. The morphology of the molten pool on the surface of the sample is more significant under low LLED, and there is a clear groove between adjacent scanning paths, which is a critical cause for diminishing the quality of powder spreading and will further damage the quality of PBF-LB formation, as shown in Figure 2(a). In short, the total surface topography is rougher since the input LLED is 0.15 J mm^–1. The surface morphology of the sample tends to be flat as the LLED is increased to 0.18 J mm^–1 and the noticeable groove phenomena between consecutive scanning pathways fade, although powder adhesion remains.

Figure 3 shows the surface roughness and relative density of Ti–6Al–4V under different LLED processes. The surface roughness of the sample is observed to be decreasing, which is mostly owing to the low LLED. And low LLED limits the fusion area between neighbouring melt channels, causing serious powder sticking and spheroidisation along the surface of the powder bed. In addition, the clear groove between adjacent scanning paths that can be seen from the SEM morphology of the untreated surface of Ti–6Al–4V above, are key factors for increasing the surface roughness of the specimen. Furthermore, it can be found that the relative density change of Ti–6Al–4V is negatively correlated with the surface roughness, which is related to the layer-by-layer powder laying and melting characteristics of the PBF-LB process, and the uneven powder bed will also increase the internal porosity. In summary, when the laser linear energy density reaches 0.21 J mm^–1, the surface roughness of the Ti–6Al–4V sample can be reduced to less than 6 μm and the relative density can reach more than 99.9%.

3.2. Microstructure analysis

Because Ti–6Al–4V is extremely sensitive to manufacturing temperature, especially in the PBF-LB process, the powder is melted into a metal liquid, causing multi-dimensional energy conduction under the action of evaporation, thermal convection, heat conduction and radiation, which rapidly lowers the temperature and shortens the solidification time, making the microstructure of PBF-LB processed samples difficult to accurately control. Figure 4 shows the microstructure of samples, which is perpendicular to the printing direction under different LLED. Inside the PBF-LB processed sample, there are a few flaws such as pores, and the microstructure is fine, consisting of ‘pin-shaped’ α/α'-Ti and irregular β grain boundaries. Figure 3(a-1 to e-1) corresponds to the magnified images of (a)–(e), which can finely portray the shape and size of the microstructure. Furthermore, the size of the microstructure becomes coarser with the increase of LLED. Obviously, at low LLED, the formation of the microstructure is insufficient and irregular due to the short period of the molten pool created by the powder under the action of the laser. The microstructure is fully grown as the input laser energy increases and the grown microstructure becomes fine and homogeneous. As shown in Figure 3(c), the LLED is 0.21 J mm^–1.

Figure 5 shows the XRD patterns of Ti–6Al–4V samples printed by PBF-LB under various LLED. The phase composition and transformation of the sample under different LLED can be properly determined by comparing the XRD standard card and according to the independent diffraction peaks and the central angle in the diffraction patterns. Observing the XRD pattern in Figure 4, under different LLED, Ti–6Al–4V samples only show multi-angle α and α’ phase diffraction peaks, which are (100), (002), (101), (102), (110), (103), (112) and (201) crystal plane diffraction peaks. However, there is no diffraction peak in the β phase. This could be since the phase content is very low and the crystal grains are exceedingly fine and distributed between grain boundaries. It is difficult to form a significant β phase grain boundary index diffraction peak.

3.3. Micro-hardness

Figure 6 shows the micro-hardness of the sample under different LLED. When LLED is 0.15, 0.18, 0.21, 0.24 and 0.27 J mm^–1, the micro-hardness values of the corresponding samples are 355.77HV_0.2, 371.43 HV_0.2, 388.17 HV_0.2, 363.51 HV_0.2 and 337.59 HV_0.2, respectively. It has been discovered that as the input laser energy increases, the micro-hardness value of PBF-LB Ti–6Al–4V samples initially increases and subsequently falls. When the LLED is 0.21 J mm^–1 with the laser power at 210 W and the scanning speed at 1000 mm s^–1, the average microhardness value of the sample S3 reaches the maximum value. When the LLED reaches to 0.27 J mm^–1, the average microhardness value of sample S5 drops significantly and the comparison of sample S3 drops by 13.03%. Furthermore, compared to S1 and S2, the micro-hardness error value distribution of samples S3, S4 and S5 is very consistent, with a moderate fluctuation value. This is mostly due to insufficient laser energy input at low line energy density, resulting in an unstable and poor continuity of molten pool, and un-melted powder or incomplete melting of the powder will restrict the fluidity of the Ti–6Al–4V solution, even preventing gas from escaping. This is the main reason for the unstable micro-hardness of samples under low line energy density.

3.4. Tensile property

The stress–strain curve of Ti–6Al–4V under tensile loading at room temperature is shown in Figure 7. The results reveal that at a loading rate of 0.5 mm min^–1, the tensile stress increases dramatically as the strain increases, and the tensile specimen enters the elastic deformation stage. When the strain rate exceeds 2%, the engineering stress area is stable and the tensile specimen is in the plastic deformation stage. In addition, under different LLED, the ultimate tensile stress of the Ti–6Al–4V specimen is between 1100 and 1200 MPa, and the ultimate strain is between 6.2% and 7.9%. In summary, the PBF-LB Ti–6Al–4V has excellent tensile properties. The tensile properties of samples S1 and S5 are far worse than S2, S3 and S4. Therefore, the laser line energy density between 0.18 and 0.24 J mm^–1 is the suitable process window for Ti–6Al–4V with high performance. At the same time, it can also be explained that low laser line energy density or high laser line energy density will have an adverse effect on PBF-LB Ti–6AL–4V. The mechanical properties of PBF-LB fabricated Ti–6Al–4V are recorded in Table 3.

Table 2

Element composition of Ti–6Al–4V powder.

Composition	Al	V	Fe	H	O	N	C	Ti
Content (wt-%)	6.14	3.87	0.21	0.0024	0.125	0.017	0.0094	Bal.

Table 3

Mechanical properties of samples with LLED of S1 0.15 J mm^–1, S2 0.18 J mm^–1, S3 0.21 J mm^–1, S4 0.24 J mm^–1 and S5 0.27 J mm^–1.

Number	UTS (MPa)	YS (MPa)	E (GPa)	Elongation (%)
S1	1180.3	1045.4	105.2	6.2
S2	1197.2	1018.5	101.1	7.5
S3	1197.5	1045.6	108.1	7.9
S4	1189.1	1037.5	100.1	7.1
S5	1162.1	1010.5	110.6	7.1

SEM was used to study the tensile fracture morphology of samples to further explore the fracture mechanism of samples under varied LLED. The results showed that the fracture is characterised by un-melted powders, micro-pore, dimples, cleavage and quasi-cleavage. Figure 8(a) shows that under low LLED, there is un-melted powder on the sample's tensile fracture surface, and the fracture is relatively flat with shallow dimples. It is an undeniable factor in reducing the tensile performance of Ti–6Al–4V. Figure 8(b) shows a smooth microporous surface that lacks the properties of ductile fracture, which can be regarded as a microporous defect. With the increment of LLED, the surface of the fracture tends to be uneven, and large and deep dimples appear. What is more, it can be seen from the schematic diagram of the partially enlarged view (c-1) that a quasi-cleavage surface appears on the surface of the fracture. When the energy density reaches 0.27 J mm^–1, the quasi-cleavage surface changes into a cleavage surface. Thus, sample S5 exhibits a mixed mode of ductile fracture and brittle fracture in Figure 8(e).

3.5. Tribological performance

Figure 9 shows the tribological properties of the samples in various LLED. Figure 9(a) depicts the time-dependent friction coefficient of the five sets of samples. When compared to the samples S3, S4 and S5, the real-time friction coefficients of S1 and S2 fluctuate substantially. The fundamental reason for this is the true contact area between the grinding ball and the micro-protrusions on the sample's surface. The S1 and S2 samples have relatively high surface roughness values (9.2 and 7.9 μm), resulting in a tiny real contact area with high stress and severe wear. It is a key factor that cannot be ignored in real-time friction coefficient fluctuations. As the LLED rises to 0.21∼0.24 J mm^–1, the surface quality of the sample is greatly improved, and the real-time friction coefficient is relatively stable. However, as the LLED reaches 0.27 J mm^–1, the real-time friction coefficient begins to fluctuate significantly, which may be caused by the ‘over-burning’ phenomenon due to the unduly high input laser energy. Furthermore, the sample's real-time friction coefficient curve shows that after around 400s of operating, the friction coefficient tends to be stable, which is also known as the stable wear stage. This is because the ‘passivation layer’ generated on the surface of the sample is gradually consumed as wear duration increases, resulting in a greater contact area and lower surface contact stress. Thus, the wear tends to be relatively stable. The average friction coefficient value is obtained according to the real-time friction coefficient value, which is shown in Figure 9(b). It can be found that in aviation lubricating oil medium, the friction coefficient of samples is between 0.145 and 0.120, showing excellent wear resistance. The order of the average friction coefficient value of samples under different LLED is S1 > S2 > S5 > S3 > S4. The average friction coefficient decreases first and then increases with the rise of LLED. The friction coefficient of S4 is 0.1505, which is reduced by 21.74% compared to S1. Figure 9(c, c-1 and c-2) show the wear three-dimensional contour surface, grinding length and width of S4, for calculating wear volume. Figure 9(d) illustrated the volume rate. It can be found that the wear rate of S1 is the highest among all samples, as high as 10.45*10⁻⁸ mm²*N⁻¹, which is inseparable from its rough surface and unstable friction coefficient. With the rise of LLED, the volume wear rate gradually drops. When the LLED is 0.24 J mm^–1, the volume wear rate drops to a minimum of 6.95*10⁻⁸ mm²*N⁻¹, which is 33.49% lower than that of S1. It can also be found that the changing trend of volume wear rate is consistent with the average friction coefficient.

Figure 10 shows the SEM wear morphology of Ti–6Al–4V samples under different LLED. Figure 10(a, b) illustrate that under low LLED, obvious furrows and severe fatigue spalling appear on the worn surface, but no loose wear debris is seen. As the LLED rises to 0.21–0.24 J mm^–1, as shown in Figure 10(c, d), the furrow wear is significantly reduced, which is replaced by adhesive wear. In addition, a small amount of wear debris can also be observed on the wear surface. As the laser line energy further increased to 0.27 J mm^–1, severe furrow wear appeared and adhesion wear became more serious. In summary, as the line energy density of the laser increases, the wear mechanism of Ti–6Al–4V formed by PBF-LB under the aviation lubricant medium first changes from severe furrow wear and fatigue wear to slight furrow wear and adhesion wear mechanism and then changes to the severe furrow wear and adhesion wear.

4. Conclusions

In this work, Ti–6Al–4V samples were successfully manufactured by PBF-LB, and the effect of LLED on the mechanical and tribological properties of PBF-LB samples was studied. The research conclusions are as follows:

With increasing LLED, the surface morphology of PBF-LB Ti–6Al–4V samples becomes flat and minimum surface roughness and maximum relative density are 5.2 μm and 99.5%, respectively. Furthermore, the microstructure consists of ‘pin-shaped’ α/α'-Ti and irregular β grain boundaries, the XRD pattern reveals that the LLED does not affect the phase composition of Ti–6Al–4V.

Micro-hardness and tensile characteristics increase and eventually decline as LLED levels rise. The maximum micro-hardness, ultimate tensile strength and maximum elongation are 388.17 HV_0.2, 1197.5 MPa and 7.9%, respectively. The tensile fracture morphology has the characteristics of un-melted powder, micro-pores, dimples, cleavage and quasi-cleavage.

When the LLED is 0.24 J mm^–1, the Ti–6Al–4V sample shows the best tribological performance and the average value of friction coefficient is only 0.1505, which is reduced by 21.74% compared to S1 (the sample printed at 0.15 J mm^–1). At this time, the volume wear rate is 6.95*10⁻⁸ mm²*N⁻¹, which is 33.49% lower than that of S1.

Footnotes

Disclosure statement

No potential conflict of interest was reported by the author(s).

Xiaojie Shi , PhD student, research direction: laser additive manufacturing.

Peipei Lu , PhD student, research direction: metal porous materials.

Xiu Ye , PhD student, research direction: laser additive manufacturing.

Shuai Ren , PhD student, research direction: numerical simulation of additive manufacturing process.

Yiyao Wang , PhD student, research direction: 3D printing and coatings.

Ziwen Xie , PhD student, research direction: laser additive manufacturing.

Yiqing Ma , PhD student, research direction: laser additive manufacturing.

Xiaojin Miao holds a PhD from Jiangnan University, lecturer, research direction: 3D printing and coatings.

Meiping Wu holds a PhD from Nanjing University of Aeronautics and Astronautics, Professor, research direction: laser additive manufacturing.

References

Kaschel

, Vijayaraghavan

, Shmeliov

, . Mechanism of stress relaxation and phase transformation in additively manufactured Ti-6Al-4V via in situ high temperature XRD and TEM analyses. Acta Mater. 2020;188:720–732.

, Guo

Effect of different lubricants on microstructural and tribological properties of TC21 titanium alloy against Si3N4 under fretting-reciprocating sliding. J Alloys Compd. 2018;743:576–585.

, Zhou

, Wang

, . Tribological investigation of additive manufacturing medical Ti6Al4V alloys against Al2O3 ceramic balls in artificial saliva. J Mech Behav Biomed Mater. 2020;104:103602.

Zhang

, Pan

, Sun

, . A net-shape forming process of Ti-6Al-4V sphere joints. Powder Metall. 2021;64:404–411.

Karayel

, Bozkurt

Additive manufacturing method and different welding applications. J Mater Res Technol. 2020;9:11424–11438.

Duan

, Zhan

, Wu

, . Analysis of elements non-uniform distribution of FeCoCrNi high-entropy alloy coatings on Ti-6Al-4V surface by laser cladding. Met Mater Int. 2021;27:467–480.

Zhou

, Li

, Yang

, . Pores structure change induced by heat treatment in cold-sprayed Ti6Al4V coating. Therm Spray Technol. 2019;28:1199–1211.

Liu

, Meng

, Zhu

, . Dynamic compressive properties and underlying failure mechanisms of selective laser melted Ti-6Al-4V alloy under high temperature and strain rate conditions. Addit Manuf. 2022;54:102772.

Cao

, Zhang

, Ryder

, . A review of the fatigue properties of additively manufactured Ti-6Al-4V. JOM. 2018;70:349–357.

10.

, Wu

, Liu

, . Study on corrosion resistance and bio-tribological behavior of porous structure based on the SLM manufactured medical Ti6Al4V. Met Mater Int. 2020;26:1182–1191. DOI:10.1007/s12540-019-00506-w.

11.

Meng

, Gao

, Yue

, . Stress-based fatigue behavior of Ti-6Al-4V alloy with a discontinuous lamellar microstructure fabricated by thermomechanical powder consolidation. Mater Sci Eng A. 2020;798:140085.

12.

Zhu

, Guo

, Ruan

, . Formation of adiabatic shear band within Ti-6Al-4V: An in-situ study with high-speed photography and temperature measurement. Int J Mech Sci. 2020;171:105401.

13.

, Lu

, Xu

, . High-performance integrated additive manufacturing with laser shock peening –induced microstructural evolution and improvement in mechanical properties of Ti6Al4V alloy components. Int J Mach Tools Manuf. 2020;148:103475.

14.

Kumar

, Prakash

, Ramamurty

Micro-and meso-structures and their influence on mechanical properties of selectively laser melted Ti-6Al-4V. Acta Mater. 2018;154:246–260.

15.

, Yu

, Wang

, . Controlling the tensile and fatigue properties of selective laser melted Ti-6Al-4V alloy by post treatment. J Alloys Compd. 2021;857:157552.

16.

Zhao

, Koizumi

, Aoyagi

, . Molten pool behavior and effect of fluid flow on solidification conditions in selective electron beam melting (SEBM) of a biomedical Co-Cr-Mo alloy. Addit Manuf. 2019;26:202–214.

17.

Sun

, Chen

, Wei

, . Microstructure and properties of high power-SLM 24CrNiMoY alloy steel at different laser energy density and tempering temperature. Powder Metall. 2020;64:1–12.

18.

Attar

, Ehtemam-Haghighi

, Kent

, . Comparative study of commercially pure titanium produced by laser engineered net shaping, selective laser melting and casting processes. Mater Sci Eng A. 2017;705:385–393.

19.

Bartolomeu

, Buciumeanu

, Pinto

, . Wear behavior of Ti6Al4V biomedical alloys processed by selective laser melting, hot pressing and conventional casting. Trans Nonferrous Metals Soc China. 2017;27:829–838.

20.

Khorasani

, Gibson

, Awan

, . The effect of SLM process parameters on density, hardness, tensile strength and surface quality of Ti-6Al-4V. Addit Manuf. 2019;25:176–186.

21.

Chen

, Zhang

, Zhou

, . Microstructural, electrochemical and wear-corrosion characterization of TI-6AL-4V-5Cu alloy fabricated by selective laser melting. J Alloys Compd. 2021;874:159953.

22.

, Ma

, Zhang

, . Effects of build direction on thermal exposure and creep performance of SLM Ti6Al4V titanium alloy. Eng Fail Anal. 2022;135:106063.

23.

Park

, Kang

, Oinam

, . Measurement of skeletal density and porosity of construction materials using a new proposed vacuum pycnometer. Measurement (Mahwah N J). 2022;196:111209.

24.

, Wu

, Liu

, . Synergistic enhancement of Re/B4C on cobalt-based composite coatings: evolution of microstructure, microhardness and tribological properties. Surf Coat Technol. 2021;425:127663.

25.

Dusza

, Csanádi

, Medveď DR , . Nanoindentation and tribology of a (Hf-Ta-Zr-Nb-Ti)C high-entropy carbide. J Eur Ceram Soc. 2021;41:5417–5426.

Study of mechanical and tribological properties of Ti–6Al–4V alloy fabricated by powder bed fusion laser beam

Abstract

Keywords

1. Introduction

2. Materials and methods

2.1. Specimens preparation

(2) where is the relative density of the sample, is water density, is the weight of sample suspended and immersed in water, is the weight of samples in air. 2.3. Mechanical tests

(3) (4) (5) where is the wear rate; is the wear volume; is the sliding distance; is the applied load; is the radius of the Si3N4 ball; is the sliding speed; is the testing time; is the wear length and is the wear width. 3. Results and discussions

3.1. Surface morphology and porosity

Footnotes

Disclosure statement

References

(2) where is the relative density of the sample, is water density, is the weight of sample suspended and immersed in water, is the weight of samples in air.
2.3. Mechanical tests

(3) (4) (5)

where is the wear rate; is the wear volume; is the sliding distance; is the applied load; is the radius of the Si₃N₄ ball; is the sliding speed; is the testing time; is the wear length and is the wear width.
3. Results and discussions