**E**ff**ects of Ultrasonic Impact Treatment on the Stress-Controlled Fatigue Performance of Additively Manufactured DMLS Ti-6Al-4V Alloy**

#### **Peter Walker 1, Sinah Malz 2, Eric Trudel 1, Shaza Nosir 1, Mostafa S.A. ElSayed 1,\* and Leo Kok <sup>3</sup>**


Received: 10 September 2019; Accepted: 18 October 2019; Published: 8 November 2019

**Abstract:** Additive manufacturing (AM) offers many advantages for the mechanical design of metal components. However, the benefits of AM are offset to a certain extent by the poor surface finish and high residual stresses resulting from the printing process, which consequently compromise the mechanical properties of the parts, particularly their fatigue performance. Ultrasonic impact treatment (UIT) is a surface modification process which is often used to increase the fatigue life of welds in ship hulls and steel bridges. This paper studies the effect of UIT on the fatigue life of Ti-6Al-4V manufactured by Direct Metal Laser Sintering (DMLS). The surface properties before and after the UIT are characterized by surface porosity, roughness, hardness and residual stresses. Results show that UIT enhances the fatigue life of DMLS Ti-6Al-4V parts by suppressing the surface defects originating from the DMLS process and inducing compressive residual stresses at the surface. At the adopted UIT application parameters, the treatment improved the fatigue performance by 200%, significantly decreased surface porosity, reduced the surface roughness by 69%, and imposed a compressive hydrostatic stress of 1644 MPa at the surface.

**Keywords:** fatigue life improvement; materials characterization; additive manufacturing; ultrasonic impact treatment; DMLS

#### **1. Introduction**

Additive Manufacturing (AM) offers great promise to the medical [1], aerospace [2], automotive and defense fields [3–5]. It provides the advantage of building complex geometries by fabricating 3D objects one layer at a time using rendered CAD models [6], as a result, several techniques of AM have been developed, including, Electron Beam Melting (EBM) [7], Selective Laser Sintering (SLS) [8], Selective Laser Melting (SLM) [9], and Direct Metal Laser Sintering (DMLS) [10]. Another advantage of AM lies within the wide range of materials that can be manufactured such as plastics [11], metals [3,12–14], ceramics [15], concrete [16] and fiber reinforced polymers [17], among others. The AM of metals are of particular interest to the production of dental implants, aerospace components, and automotive structures [18]. Today, stainless steel [19,20], nickel alloys [21,22], aluminum alloys [23,24] and titanium alloys [25,26] are common materials for metals AM.

DMLS is one of the most common AM processes for 3D printed metals because it maintains dimensional control while producing complex features at high resolution [3]. There are two primary

methods of manufacturing by DMLS: powder bed and powder deposition [27]. Powder bed methods rely on a high energy source, typically a laser (although some similar systems use electron beams), to locally sinter or melt metal particles on a powder platform. New powder layers are periodically added while the platform is levelled to accommodate the addition of new material. The 3D model is constructed in a single vertical direction [6]. Alternatively, powder deposition directly deposits the metal powder and melts it in place using a high-powered laser [28]. Unlike the powder bed method, which is typically restricted to one type of alloy, powder deposition has the ability to include different metal powders for functionally graded materials [29].

This process however, also has a number of drawbacks. Parts produced via DMLS typically have poorer mechanical properties compared to those produced by traditional means, which has relegated the potential uses to prototypes and short-term tooling operations [30]. Incomplete powder melting often leads to very rough surface finish and porosity, which in addition to being aesthetically displeasing also compromises fatigue life and can be a significant issue for wetted surfaces of air and water craft [29,31]. In addition, the rapid heating and effective quenching of the metal results in a highly martensitic microstructure in most alloys [32]. While this produces a material with a high yield strength, it is also intensely brittle [32]. Though the microstructure resists the formation of cracks, once the cracks themselves form, the propagation of the cracks is quite rapid [33]. The high temperature gradient involved in these processes frequently causes thermal stress that compromise the fatigue performance of metals produced by DMLS techniques.

Titanium alloys such as Ti-6Al-4V are commercially available for additive manufacturing. Due to its high strength to weight ratio and fracture toughness, it is an ideal alloy for a wide range of applications in the aerospace and biomedical engineering fields [34]. For instance, Ti-6Al-4V alloy is used in dental laboratories for medical implants and prosthetics due to its corrosion resistance, high specific strength as well as its biocompatibility characteristics. AM also allows for the creation of porous titanium structures that help facilitating bone ingrowth and adhesion for implants [35]. Advances in topology optimization allows for hyper-efficient geometries to be produced exclusively through additive manufacturing. One notable example includes the Airbus A320 nacelle hinge bracket which was eventually produced by AM [36]. With the many advancements in the field of AM, the production of titanium parts has become economically viable. Hence it is imperative that the fatigue properties are improved for the next generation of additively manufactured titanium components [33,37–40]. In the case of Ti-6Al-4V, the hexagonal close-packed (HCP) α phase and trace amounts of a body centered cubic (BCC) β phase are almost entirely replaced with the martensitic α' phase. The poor surface finish and high porosity are also factors that offset the microstructural characteristics on the fatigue performance.

The demand for functional AM parts has been rising as high reduction in assembly costs are available coupled with decreases in mass. One example is the fuel nozzles for the General Electric (GE) LEAP engine have been additively manufactured to be 25% lighter while eliminating previous models that required laborious assembly. Its successful design has now been 3D printed more than 30,000 times since its conception [41]. With the increasing demand for metal AM, researchers have begun to develop techniques to improve the fatigue performance by improving the surface finish and by inducing compressive residual stresses at the surface. Shot peening [42], Ultrasonic Nanocrystal Surface Modification (UNSM) [43–46] and grit blasting [47] are well known examples of beneficial treatments on AM of metals. However, none of these processes were able to address all the following simultaneously: surface roughness, surface porosity, fatigue life and tensile residual stresses.

Studies have shown that Ultrasonic Impact Treatment (UIT) improves the surface finish and fatigue properties in the field of post-welding [48]. UIT is a process in which an indenter vibrating at ultrasonic frequencies slides over a surface. This treatment plastically deforms the surface, improving the surface finish while inducing and redistributing residual stress in the part resulting in enhanced fatigue life [48]. UIT devices operate by inducing plastic deformation from the indenter or impact needle by first exciting a transducer by a controlled voltage input. The power source directly controls the oscillations exhibited by the transducer, which sends its frequencies to the sonotrode (ultrasonic horn) [48]. The research of E. Statnikov et al. [49] compared a variety of methods that improve the fatigue life of welded joints. Other similar methods include Hammer Peening, Shot Peening and Tungsten Inert Gas (TIG) dressing. An improvement of 65% was observed in the UIT joints. In B.N. Mordyuk et al. [50] investigated the enhancements that occur in the surface layer of ultrasonic impacted specimens. It was concluded that the compressive residual stresses and work hardening of the surface layer attributed to the improvement in fatigue properties of processed specimens. A.I. Dekhtyar et al. [51] concluded that at high stress levels, UIT-processed Ti-6Al-4V has a fatigue life that is twice of the pristine (untreated) samples, and a roughness reduction Ra of 75%. While UIT is widely used in fatigue improvement of welded joints, the aim of this paper is to treat the surface of Ti-6Al-4V specimens produced by DMLS using UIT to improve roughness, increase hardness and induce surface compressive residual stresses to enhance the fatigue life of the components.

It is well known that the resistance to tensile fatigue of a metal will increase by the addition of compressive stress [37]. UIT compares favorably to other surface treatments due to its higher impact forces [48]. Quantifying this level of stress by the cold working of the surface for titanium alloys can prove to be beneficial in future engineering applications, as evaluation of impact forces between the pins and the metal surface of UIT is under researched. Force estimations for other cold hardening treatments such as shot peening have been measured. One method used acoustic emissions for velocities of 30 to 88 m/s; however, this was not able to properly determine the impact force [52]. Many other techniques have been developed for measuring shot peening impact forces, which include a shot peening intensity detector; however, most methods cannot be directly applied to measure forces for an ultrasonic impact device [53]. Research into the optimization of UIT parameters has led to the use of oscilloscopes for defining the impact characteristics between the pin and the metal surface [48,54]. The frequencies were able to illustrate the ultrasonic deformations and elastic recovery of the process. The frequency patterns are also able to show the stochastic nature of the impacts that are influenced by both the impact depth and plastic deformation. An additional experiment presented in this paper is conducted to quantify impact forces during UIT on DMLS printed titanium specimens.

#### **2. Materials and Methods**

#### *2.1. Specimens Manufacturing*

Flat dog-bone fatigue specimens were manufactured in accordance to ASTM E466-15 [55] for force controlled fatigue testing. Figure 1a illustrates the dimensions of the dog-bone in millimeters. The titanium dog-bone specimens were built using Ti-6Al-4V grades 23 powder from AP&C, composed of between 5.5 and 6.75 wt% aluminum, 3.5 and 4.5 wt% vanadium and <0.25 wt% iron and trace amounts of <1 wt% impurities such as oxygen, nitrogen and hydrogen with the balance being titanium [56]. The 3D printer used to manufacture the specimens was an EOSINT M290/400W machine following a general process parameter of layer thickness of 30 μm and volume rate of 5 mm3/s where the volume rate is a measure of the build speed during laser exposure of the skin area [57]. The process parameters are optimized in such a way as to provide mechanical properties comparable to other literature [58]. The particle size ranged from 15 to 23 μm and the printing layers thickness were 60 μm. The specimens were heat treated in accordance to AMS 2801 to relieve the residual stress induced by the rapid melting and solidification that takes place during the printing process [26,59]. Figure 1b shows the printing orientation, where the printing platform lies in the x–y plane. A single 2D layer is formed on the platform when the laser beam sinters particles starting from the left end and moving towards the right end of the dog-bone shown in Figure 1.

**Figure 1.** Specimens design and preparation. (**a**) Schematic of flat sheet fatigue specimen with rectangular cross-section; (**b**) orientation of the specimens during DMLS manufacturing process.

When the layer is completed, the platform is levelled in the z-axis allowing a new layer to be added on top of the previous one. Wire Electrical Discharge Machining (EDM) is a precise cutting technique that minimizes the need for excessive post processing machining and was used to breakaway support and remove the specimens from the platform [60]. Lastly, the edges were then polished using an emery cloth of grades 120 and 220 [61].

#### *2.2. UIT Device*

The UIT device, displayed in Figure 2, is a 20k Ultrasonic Impact Treatment device, DW-CJ20-1000 produced by Dowell Ultrasonics [62], typically used as a hand-held tool for post-welding processing. It consists of a power supply, shown in Figure 2a, the UIT tool, shown in Figure 2b, the impactor head, shown in Figure 2c and the ultrasonic generator, shown in Figure 2d. It is equipped with slots for four impactors, but for the purposes of this experiment—only one impactor was used, as illustrated in Figure 2c. To provide automated control, a custom-built fixture is used to attach the device to the spindle of a Computer Numerical Control (CNC) Mill. Figure 2e shows the treatment path programmed to minimize the surface roughness while obtaining a uniformly deformed surface. The scanning speed of the CNC machine was set to 1000 mm/min. The spacing between the scans is known as an interval. Amplitude control of the device is controlled by a Fagor 8035M controller [63]. A constant amplitude of 57% of 40 μm was used during treatment as testing showed that this amplitude provided the most consistent plastic deformation of the surface without damaging the samples. The interval for treating titanium alloys typically ranges between 10 and 70 μm [43–46,64]. The path contours were chosen to increase outwards at 71.1 μm intervals to match previous efforts and research into surface treatments on titanium [43,44,65].

**Figure 2.** UIT system and application pattern. (**a**) Power supply unit; (**b**) side view of the UIT device; (**c**) front view of the UIT device showing impactor position; (**d**) ultrasonic generator; (**e**) schematic of treatment path.

#### *2.3. Specimens Fixture*

To apply the UIT a new fixture was designed, as shown in Figure 3. The Ti-6Al-4V specimens were clamped to an aluminum plate that was supported by four steel rods and four aluminum sleeves, as illustrated in Figure 3. The plate can freely slide along the rods and its motion range is limited to four high precision springs from McMaster Carr [66] placed between the plate and the end of the supporting rods. Compressing the springs allows for a constant static force to be applied onto the samples during treatment. By pushing the UIT device into the plate and compressing the springs a certain distance, the amount of static force can be determined. Aluminum cutting fluid [67] was used to lubricate the rods so that the plate could freely move by the springs. For this experiment, a static force of 30 N was applied at the tool specific constant frequency of 19.86 kHz.

**Figure 3.** Spring assembly fixture with a clamp mounted to the CNC machine.

#### *2.4. Fatigue Testing*

Force controlled fatigue tests were undertaken with a servo-hydraulic 810 Material Testing System (maximum load 100 kN) [68]. The MTS consists of an upper and a lower clamping grip. For consistent and precise alignment, a fence was attached to each grip. All specimens were tested under a clamping pressure of 5.52 MPa, frequency of 25 Hz, maximum stress level of 400 MPa and mean stress level of 200 MPa.

#### *2.5. Microscopy*

Microscopic observations were conducted using a combination of optical and scanning electronic microscopy. The former was performed with an Inverted Trinocular Metallurgical Microscope including an AmScope 18 MP MU 1803 Camera [69]. Electron microscopy was performed with a Zeiss GeminiSEM 500 at the University of Ottawa's Centre for Photonics Research [70].

#### *2.6. Roughness*

Roughness measurements were performed using DektakXT Stylus Profiler by Bruker [71]. A Peak and Valley analysis was conducted to determine the relative roughness of each sample. As illustrated in Figure 4, a lateral and a longitudinal line scan is performed on each side of a specimen. The scans intersect to form a cross. Roughness measurements were taken over 2 mm in each direction and the results are averaged.

**Figure 4.** Orientation of line scans for surface roughness.

#### *2.7. Hardness*

Rockwell hardness was used to analyze the difference in hardness between treated and untreated specimens following the testing standards of ASTM E18-18a [72]. A Rockwell C test was performed using a test force of 150 kgf on treated and untreated specimens using "The Portable Rockwell Hardness Tester" by Bowers [73].

#### *2.8. Residual Stresses by X-ray Di*ff*raction*

Residual stress is defined as the stress remaining in a solid material after an applied force or plastic deformation has taken place. UIT imposes a high plastic strain at the surface of the treated surface, which results in compressive residual stresses. X-ray diffraction (XRD) is a non-destructive method for analyzing the residual stress in a material and is the method of choice for this paper.

Chemical etching of the surface was applied to remove suspected amorphous or oxide material layers, potentially caused by initial stress relaxation treatment. The effect of etching the surface and XRD quality of the scans is illustrated in Figure 5. Peak shapes become much clearer and better defined. It is also important to note that mechanical polishing or grinding will cause lattice strains to be formed at the surface and are not recommended for cleaning samples for residual stress measurements [74]. To conserve and reveal the surface stress layer, a chemical etchant—Kroll's Reagent—was used to remove small amounts of material from the top of the treated and untreated titanium specimens [75,76]. The amount of material removed from the surface was measured to be on average ∼ 8 μm.

**Figure 5.** XRD Comparison between etching and without etching.

It is suggested by P. Mercelis et al. [40] that the surface porosity of DMLS manufactured parts poses difficulties for measuring residual stresses due to the presence of zero-stress porosity borders. Stress discontinuities on the surface result in lower residual stresses to be measured than in reality. Hence, cleaning the surface of the samples through etching would also help to reduce the effects of roughness and porosity that may negatively affect the XRD measurements.

The XRD measurements were taken using a Malvern Panalytical Emperyean [77], shown in Figure 6b. The machines power was set to 40 kV with a current of 40 mA for a strong signal response—particularly at higher rocking angles. A half degree diffraction slit was used with a mask size of 2 mm and an anti-scatter slit of 2◦. The diffraction arm was a Branson Bragg attachment with a 0.04 mm slit. The incident arm was equipped with a 0.04 mm slit and a 9.1 mm opening.

**Figure 6.** XRD System. (**a**) Drawing of diffraction directions along the surface and at angles φ and ψ. Both σ<sup>1</sup> and σ<sup>2</sup> are perpendicular and reside in the plane of the specimen surface; (**b**) sample oriented at 0◦ on multi-purpose stage of Emperyean system.

The analysis method for computing the stress tensor was the Winholtz-Cohen Least squares Analysis [78]. A total of 36 measurements per sample were used for the analysis, each representing a unique tilt and azimuth angles. The chosen x-ray elastic constants of the DMLS Ti-6Al-4V specimens where 2.355 <sup>×</sup> 10−<sup>5</sup> MPa−<sup>1</sup> as the S2 constant and <sup>−</sup>2.9877 <sup>×</sup> 10−<sup>6</sup> MPa−<sup>1</sup> for the S1 constant [65,79]. These constants are the interplanar properties for the bulk alpha phase of the metal alloy. The X-ray wavelength was set to 1.519 Å.

The specimen coordinate system had the azimuth or phi (φ) of value zero lined up with the horizontal direction (X) as illustrated by the cross seen in Figure 6a. The tilt angles or psi (ψ) is shown in Figure 6a where the 0◦ begins at 90◦ from the samples surface.

#### *2.9. Estimation of Impact Force*

A piezoelectric force sensor was used to estimate the impact forces at the surface during treatments. To measure the reaction forces at the surface; an alternative fixture was developed to house the sensor and impact plates. This fixture contains linear ball bearings instead of steel bushings to guide the plate, as seen in Figure 7a. The support rods were thickened, and the four previous compression springs were replaced with two larger springs. Proper calibration using a strain gauge was performed so that the correct amount of static force could be determined based on the compression of the spring system. The sensor was attached to the moving plate on the fixture with a titanium specimen used as an impact cap. Operation of the UIT during the estimation of impact force was controlled in much of the same way as during the UIT. The sensor would register the forces from the UIT impacts on the titanium cap. For simplistic design, the cap was a repurposed fatigue specimen.

**Figure 7.** Experimental setup for impact force estimation. (**a**–**c**) Impact force fixture; (**d**) impact areas and distances.

Due to the design of a repurposed fatigue specimen as an impact surface, reaction forces can be registered by the sensor depending on its distance from the impact. The titanium cap was treated as a double supported beam with one end constrained by a bolt and the other end constrained by the sensor using a double-sided thread. The double supported beam assumption allows for a simple ratio to be developed based equilibrium of moments to solve for the actual impact forces. A diagram of the impact set up can be seen in Figure 7b,c The selected distances between the sensor and the impact region are displayed in Figure 7d, where *d1* is 8.89 mm and *d2* is 47.752 mm. The relationship between the sensor reaction force and the impact force is presented in Equation (1).

$$F\_{\text{input}} = -F\_{\text{Sensar}} \frac{d\_2}{d\_1} \tag{1}$$

where Fimpact is the force of impact and Fsensor is the force registered by the sensor.

Data acquisition was performed using a Lecroy WaveSurfer 3000 Oscilloscope [80]. The force sensor was a PCB piezoelectric force sensor model 208C05. The force range of the sensor was 0 to 4500 kgf. The conversion from voltage to force was assumed to be linear with a sensitivity of 0.2170 mV/N (±15%).

Calibration of the sensor was checked using test procedure AT501-5. The amplifier gain of the sensor was set to 100. Additional shrink tubing was added between the sensor and the wire nuts to prevent vibrations during testing from unscrewing and ejecting the wires from the sensor. The peak voltage read by the sensor for a short ultrasonic impulse was regarded as the impact force for a given static load. The compression of the springs drives the static load with the total stiffness of the system equal to 3.06 N/mm.

#### *2.10. Microstructural Analysis*

The effect of the UIT on the microstructure of the material was investigated through Light Optical Microscopy (LOM) and Scanning Electron Microscopy (SEM), the latter being done with both secondary electron and backscatter electron (BSE) detection. Metallographic samples were cut from tensile specimens and mounted in a resin made from phenocure combined with ~15% Technotherm conductive powder for improved SEM performance [81]. Samples were polished with an Allied High Tech MetPrep 3 system with a PH-3 powerhead [82], using SiC paper with a CAMI Grit designation ranging from 120 to 1200 [82], followed by a polish with a suspension of 30% 0.3 μm alumina particles and a final polish of 30% H2O2 and 30% 0.05 μm particles [83]. Samples were etched prior to microscopy, using both classical Kroll's reagent and a 10% HF etchant for better grain boundary definition, in keeping with ASTM Standard E407-99 [84].

#### **3. Results and Discussion**

#### *3.1. Fatigue Life*

A stress-controlled fatigue test was conducted according to DIN 50100 [85] and based on a logarithmic normal distribution. Table 1 displays the fatigue life of Ti-6Al-4V specimens before and after UIT application. The specimens were tested in a servo-hydraulic 810 Material Test system (MTS) at a peak stress level of 400 MPa and minimum stress of 0 MPa. The average number of cycles for the treated and untreated specimens is compared in Figure 8.


**Table 1.** Fatigue life of treated and untreated specimens at 400 MPa.

Evaluation of the fatigue performance of DMLS Ti-6Al-4V shows that the fatigue life of untreated specimens is 77% lower than handbook values [37]. The results in Table 1 and Figure 8 show that the fatigue life of treated specimens at 400 MPa corresponds to a 200% increase compared with that of untreated specimens. In other words, the lifetime of the treated samples is three times as long as untreated samples. D. Cattoni et al. [47] only achieved a 4% fatigue improvement by blasting Ti-6Al-7Nb, whereas A.I. Dekhtyar et al. [51] prolonged the lifetime of Ti-6Al-4V, manufactured using the cost-effective blended elemental powder metallurgy technique, by two orders of magnitude after applying ultrasonic impact treatment.

**Figure 8.** Average fatigue life of treated and untreated Ti-6Al-4V specimens tested at 400 MPa.

#### *3.2. Surface Microscopy*

Figure 9a illustrates the surface of Ti-6Al-4V as manufactured by DMLS, which is dominated by the partial melted powder spheres from the additive manufacturing process along with numerous hills and valleys. Despite extensive cleaning in an ultrasonic bath with organic solvents, the rough surface still traps surface contaminants. Tracklines formed by the solidification of the powder bed layers are also visible. These further add to the roughness of the surface and can become sights of crack nucleation and potentially propagation as well [59]. Figure 9b shows the boundary zone between the treated and untreated zones, and the contrast is quite striking. The treated zone is of relatively uniform height, with the spherical powders having been plastically deformed by the impact treatment. However, even this small image of the boundary zone shows that the boundary is quite ragged, with "peninsulas" jutting out from the treated zone and small "islands" of treated area that are separated from the treated zone. The fully treated zone is shown in Figure 9c, and from simple observation it is clear that the surface is much less rough and thus much less conducive to crack nucleation.

**Figure 9.** Specimens surface microscopy. (**a**) Surface of untreated specimen; (**b**) boundary zone between treated and untreated zone; (**c**) treated surface of specimen. All images taken with SE2 detector at 15 kV.

The UIT had successfully improved the surface finish of the DMLS Ti-6Al-4V, which corresponds to the improved fatigue performance. Unlike the surface of metals treated by grit blasting and UNSM, the surface of UIT specimens is not jagged. This is due to the combination of vibrating the indenter at ultrasonic speeds while sliding across the surface.

#### *3.3. Roughness*

The results from the DektakXT profilometer used to perform a Hills and Valleys profile on both treated and untreated specimens is displayed in Figure 10. The untreated surface is 3.2 times rougher than the treated surface. A control factor for fatigue strength in specimens with high surface roughness is crack propagation, whereas for specimens with low surface roughness it is crack initiation [86]. Additional stress raisers are introduced as a result of rougher surfaces; increasing the number of potential crack initiation sites. Stress raisers as a result of rougher surfaces will reduce crack initiation life and the fatigue limit. Therefore, smooth surfaces can be considered as a contributing factor to the improved fatigue life, as seen by the fatigue performance of the treated samples with improved surface roughness.

**Figure 10.** The average roughness of treated and untreated specimens measured by DektakXT.

The effects of UIT on the AM titanium surface are also observed using a two-dimensional map scan as shown in Figure 11. The peaks and valley analysis clearly illustrate the effects on both surface roughness and surface porosity. The untreated surface has a much higher density of hills and valleys unlike the treated surface, which has a much more uniform surface and less pronounced differences between peaks and valleys. The highly irregular surface of the DMLS titanium is attributed to the manufacturing process and incomplete fusion. The final layer of metal powder during printing is not completely melted and produces a very rough surface. The high impact forces of the UIT helps diminish these effects and improves the surface quality and fatigue performance.

**Figure 11.** Roughness map scans between the (**a**) untreated and (**b**) treated surfaces.

#### *3.4. Hardness*

The results from Rockwell C tests performed using a test force of 150 kgf are illustrated in Figure 12, the treated specimens are on average 21% harder than the untreated specimens. However, it is important to note that the hardness is highly variable in both samples. The lowest hardness values from the treated sample are roughly equivalent to the hardest values from the untreated sample. This variability is likely due to the very rough surfaces of both samples. As a result, no statistical conclusion can be drawn from the hardness data about which sample is harder overall. It can be said that the hardness of the treated specimen is more consistent than that of the untreated specimen, and that there are points on the treated specimen that are harder than points on the untreated specimens. This possible increase in hardness at certain points could be attributed to the work hardening done by the UIT, a result of the increase in dislocations density at the surface which will resist further deformation. The increase in hardening could lead to a resistance to crack initiation sites in those locations, which corresponds to an increase in fatigue life [65].

**Figure 12.** The average hardness of treated and untreated specimens from the Rockwell C test.

#### *3.5. Residual Stresses*

For values ranging between 2.5 and 5 μm in depth, residual stresses play a predominant role in improving fatigue life [51]. Based on the results of the residual stress and residual strain measurements, shown in Tables 2 and 3, respectively, the success of the UIT can be attributed to high compressive stresses at the surface of the specimens.




**Table 3.** Residual strains results.

These stresses help minimize the generation of tensile stresses at the surface due to cyclic loadings thereby suppressing crack formation and nucleation [49].

The high compressive stresses help minimize the damage caused during cyclic loading by shifting the mean stress downwards. This eventually leads to longer life in the part as compressive stresses will lower the stress ratio. Experimental data has shown that as the stress ratio becomes increasingly negative, longer lives were measured. This can be explained by compressive surface stresses preventing dislocations from moving within the material. The hydrostatic stresses for each tensor reveal the same trend; the untreated sample had a hydrostatic stress of −265 MPa while the treated sample had −1644 MPa. A high compressive hydrostatic stress could also represent an increase in fatigue resistance due to UIT.

#### *3.6. Impact Force Quantification*

The impact force with respect to the variation of compressive static force was obtained. The results show that for a static force between 0 and 10 N, the impact force is relatively constant and is on average 20 kN. After which, the force steadily increases up to 72 kN when the static force is equal to 30 N. The trend is slightly parabolic and impact forces at zero newton of static force were measured to be slightly higher than at 10 N of static force. The impact force on the fatigue samples is thus 72 ± 11 kN. For every second, approximately 2.1 <sup>×</sup> 104 <sup>±</sup> 6.0 <sup>×</sup> 102 impacts occur at a static force of near zero, seen in Figure 13. This would mean that for an amplitude of 57% of 40 μm the average absolute vertical speed of the indenter is 0.94 × 0.03 m/s.

**Figure 13.** Force vs. time as measured by the force sensor.

The high impact forces are due to metal on metal deformation between the steel pin indenter and the titanium surface. High forces are registered due to the high modulus of elasticity of the mediums and by the ultrasonic vibrations and stress waves [48]. Another factor to consider for high impact forces are due to impulses occurring over a very short time period. The force versus time plots in Figure 13 show the readings of the oscilloscope of a low static force. When minimal static forces are present, the force pattern is rather predictable and consistent. As the static force is increased, the level of stochastic behavior increases dramatically, as seen in Figure 14. The increased randomness is attributed to the increase in plastic deformation which creates widening gaps with varying depth. The subsequent plastic deformation on the surface will then begin to slightly alter the impact forces slightly due to both work-hardening and rebounding forces from the support springs.

**Figure 14.** Impact force (kN) vs. static load (N).

The impact force on the surface, given one UIT needle, can be replicated if an equivalent force is used to generate the plastic deformation. Hence, other work hardening processes can achieve the same effects of the UIT if similar impact forces are used combined with a transverse velocity along the surface. Possible errors in the impact force assessment can be attributed to off-axis impact forces. Loads applied to the side of the sensor may cause higher than normal readings. This is a result of coupling forces acting on the sensor. However, calibration between the distances of the applied load to sensor was made to ensure that off-axis effects are diminished.

The saturation point of the sensor was reached once the static forces approached 30 N as it was close to the maximum range of the sensor. Any other force impact reading after the saturation point would not have been reliable as the sensor would have been destroyed. It is unknown what impact forces are possible after 30 N as extrapolation of the data is within an area of high uncertainty. The results are also calibrated for the UIT fixtures set up which includes a spring-loaded system. The springs vibrations combined with the indenter's oscillations creates a varied and stochastic process leading to large error bounds during the treatment process. The reaction forces of the springs may also contribute to higher than normal impact forces. Due to the calibration of the sensor in combination with the fixture, there is a level of specificity in the force calibration curve that may limit its applicability for future engineering applications.

#### *3.7. Microstructure Analysis*

As mentioned previously, the rapid heating and cooling cycle created by most forms of additive manufacturing processes results in a heavily martensitic microstructure. In the case of Ti-6Al-4V, this is characterized by a so called "basket-weave" pattern. The comparison between the microstructure of conventionally wrought Ti-6Al-4V, shown in Figure 15a, and the microstructure of additively manufactured Ti-6Al-4V, shown in Figure 15b, is quite striking. The basket-weave pattern of the

overlapping α' phase is quite visible, compared to the more randomly orientated microstructure produced by more conventional means.

**Figure 15.** Optical micrographs of (**a**) Wrought Ti-6Al-4V and (**b**) DMLS Ti-6Al-4V.

Neither optical microscopy nor electron microscopy of the treated vs. untreated samples showed a significant change in microstructure. In both cases, the distribution of the α' and β phases appear randomly distributed, with no obvious relationship between the β phase and the surface of the specimens, as seen in Figure 16a,b. Due to the low temperatures observed during the cold working (the samples were too hot to touch immediately after processing, but did not show any evidence of oxidization) the samples never reached a point where dynamic recrystallization could occur, so microstructural changes would be unlikely.

**Figure 16.** SEM micrographs of (**a**) UIT applied and (**b**) untreated specimens, BSE at 30 kV.

In both cases β phase "needles" were observed at various angles to the surface of the sample, and their lengths were relatively consistent with an average 17.2 μm at the surface in the untreated region and an average 16.8 μm at the surface of the treated region, a 2.3% difference. This suggests that the predominant mechanisms of fatigue life improvement are the addition of compressive residual stress and surface roughness improvements.

#### **4. Conclusions**

By applying UIT to the surface of DMLS Ti-6Al-4V, a nearly three-fold increase in fatigue life for cyclic stress levels of 400 MPa was achieved. This improvement can be attributed to the several aspects of the treatment, reduction in surface porosity, a decrease in surface roughness and possibly causing local increases in hardness. By cold working the surface, the barrier for crack nucleation is increased and delaying the onset of fatigue cracking at the surface. The large compressive stresses

imposed by UIT on the treated area also suppress crack propagation by offsetting tensile stresses at the surface during cyclic loads. In addition, improved surface finish due to UIT can help to reduce the number of potential sites for crack nucleation by reducing surface porosity, again delaying the onset of crack formation. The impact force of the treatment was determined to be 72 kN and a curve of impact force versus compressive static force was developed. Microstructural observations did not demonstrate any significant change between treated and untreated specimen edges, suggesting the predominant mechanisms of the treatment are the surface modifications and compressive residual stress. In conclusion, UIT can successfully improve the surface finish while simultaneously improve the fatigue life.

**Author Contributions:** Conceptualization, M.S.A.E., and L.K.; methodology, P.W., S.M., E.T. and S.N.; software, E.T.; validation, E.T., P.W.; formal analysis, E.T. and P.W.; investigation, P.W., S.M., E.T. and S.N.; resources, M.S.A.E.; data curation, E.T, P.W.; writing—original draft preparation, E.T., P.W. and S.N.; writing—review and editing, P.W. and M.S.A.E.; visualization, E.T., P.W. and S.N.; supervision, M.S.A.E.; project administration, M.S.A.E.; funding acquisition, M.S.A.E.

**Funding:** This research was funded by BOMBARDIER Aerospace of Montreal, in collaboration with CARIC National Forum (grant number CARIC-CRIAQ MDO-1601\_TRL4+) and MITACS Canada (grant number IT07461).

**Acknowledgments:** Mostafa S.A. ElSayed acknowledges the financial support provided by BOMBARDIER Aerospace of Montreal, in collaboration with CARIC National Forum (grant number CARIC-CRIAQ MDO-1601\_TRL4+) and MITACS Canada (grant number IT07461). The authors would like to extend their acknowledgements to the employees of the machine shop, and Dix of the X-ray Diffractometry Lab, both at Carleton University for their guidance and patience, the Centre of Photonics Research and Jeff Ovens at the X-ray Core Facility, both at the University of Ottawa for their training and assistance.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


© 2019 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (http://creativecommons.org/licenses/by/4.0/).

### *Review* **Part Functionality Alterations Induced by Changes of Surface Integrity in Metal Milling Process: A Review**

#### **Caixu Yue 1,\*, Haining Gao 1, Xianli Liu <sup>1</sup> and Steven Y. Liang <sup>2</sup>**


Received: 19 September 2018; Accepted: 14 November 2018; Published: 9 December 2018

**Abstract:** It has been proved that surface integrity alteration induced by machining process has a profound influence on the performance of a component. As a widely used processing technology, milling technology can process parts of different quality grades according to the processing conditions. The different cutting conditions will directly affect the surface state of the machined parts (surface texture, surface morphology, surface residual stress, etc.) and affect the final performance of the workpiece. Therefore, it is of great significance to reveal the mapping relationship between working conditions, surface integrity, and parts performance in milling process for the rational selection of cutting conditions. The effects of cutting parameters such as cutting speed, feed speed, cutting depth, and tool wear on the machined surface integrity during milling are emphatically reviewed. At the same time, the relationship between the machined surface integrity and the performance of parts is also revealed. Furthermore, problems that exist in the study of surface integrity and workpiece performance in milling process are pointed out and we also suggest that more research should be conducted in this area in future.

**Keywords:** milling process; part functionality; surface integrity; research progress

#### **1. Introduction**

In today's competitive manufacturing industry, the ultimate goal of manufacturers is to produce higher-performance products at lower cost and in less time. At the end of the machining process chain, thermomechanical coupling load has significant impact on workpiece surface property with a direct link to functionality. To get a reliable machined component with high fatigue strength, high wear resistance, and high dimensional accuracy are the goals of the machining process. It is necessary to evaluate the effects of machining parameters, tool parameters, etc. on machined surface performance capabilities.

The specified term surface integrity can be used to evaluate the machined surface properties after manufacturing operations. There are mainly some aspects to describe surface integrity: topography, metallurgy characteristics and residual stress. The topography is made up of surface roughness, waviness and flaws. The metallurgy characteristics include grain size, plastic deformation, microhardness, phase transformation, recrystallization, etc. [1].

The door to the study of machined surface integrity is opened by a review article written by Field et al. [2]. The results show that surface integrity is the intrinsic property of surface hardening conditions that are produced after processing [3]. Then, considerable research on machined surface integrity was carried out [4–8]. The CIRP (The International Academy for Production Engineering) published its keynote paper "Capability Profile of Hard Cutting and Grinding Processes" in 2005 to guide the research of surface integrity; subsequently, a collaborative working group on surface integrity and functional performance of components was established in 2008 [9].

The ultimate research goal of surface integrity is to get the component with high performance capacity. In the cutting process, the finishing pass will define the thermos and mechanical state of the machined surface. Surface integrity has a significant impact on several relevant characteristics of the final functionality of the component, such as dimensional accuracy, friction coefficient, wear and thermal resistance, and fatigue behavior corrosion, as shown in Figure 1. For example, the finished component may lose efficacy for many reasons, such as changes in dimensions due to wear or plastic deformation, deterioration of the surface finish, and cracking or breakage [10]. Therefore, it is important to reveal the effect of the manufacturing process on finish part functionality.

**Figure 1.** The effect of surface integrity on part functionality.

The definition of surface function differs depending on the operating performance of the surface. For example, when translational surfaces are used as surfaces for the bearing inner and outer rings, tribological functionality will become dominant. However, for the applications of dynamic loading and cyclical stresses, fatigue characteristics may be a prime consideration in determining component failure [11]. Therefore, in order to get the influence mechanism of cutting parameters on parts performance, it is important to study the intrinsic relationship between the factors of surface integrity and parts performance [6].

Milling has the advantages of high production efficiency and wide processing range, and it is widely used in key fields such as aerospace, mold, automobile, and parts manufacturing [12]. To reveal the topic more clearly, the surface integrity and functionality of the part after the milling process are focused on in this paper. The effects of the manufacture process on surface integrity are discussed. Through the analysis research and the finite element modeling method, the surface integrity can be better understood. Then, a review of the state-of-art research on finished component functionality affected by surface integrity, such as surface topography, surface metallography, and residual stress, is presented. In this review, the component functionality mainly includes fatigue strength and wear resistance.

#### **2. Part Functionality as Effected by Surface Topography**

#### *2.1. Machined Surface Topography Characteristic*

The surface topography of any manufacturing process is a critical index. It is mainly affected by geometric properties of the machining system for milling process. Choosing suitable cutting parameters for given tools, workpiece materials, and machine tools is an important step to get reasonable surface

topography. Because the two-dimensional parameters cannot properly represent the surface features of milling, more and more attention is being paid to three-dimensional parameters today.

Considerable research has been performed to get suitable surface topography. Zhang et al. [13] provided in-depth characterization and analysis of the three-dimensional topography of feed direction and cross-feed direction in the hard milling process of AISI H13 steel (AISI is an abbreviation of American Iron and Steel Institute), as shown in the Figure 2. The results show that better surface morphology can be obtained by using higher cutting speed, lower feed speed, and lower cutting depth. Wang et al. [14] studied the surface morphology and surface roughness in the milling process of AlMn1Cu. Due to the high ductility of the AlMn1Cu material, the material flows plastically along the side-cutting edge and meanwhile is extruded by the side-cutting edge to cause the material to accumulate on the machined surface. With the increase of cutting depth and feed per tooth, the plastic flow of the material along the cutting edge is strengthened, resulting in increased surface roughness. However, with the increase of cutting speed, the plastic flow of the material is weakened along the secondary cutting edge, resulting in that the surface roughness value decreases. Ghani et al. [15] found that high cutting speed, low feed rate, and low cutting depth can be used to obtain ideal surface quality in the semifinishing and finishing process of AISI H13 steel.

**Figure 2.** Surface topography of milling surface [13]. (**a**) Surface texture zones in milling and (**b**) surface topography by milling process.

Some scholars carried out the research of prediction of surface topography for milling process. As established by experimental tests, the actual roughness values obtained usually deviates from theoretical ones. There are several reasons accounting for this, such as runout of the milling cutter tooth tips, rounding of the cutting edge, and irregularities of cutting edge [16].

In order to design a machining process for blades in turbine engines without prior experiment, Denkena et al. [17] developed a model to predict the surface topography. The effect of stochastic topography on the flow losses was investigated also in his research. The machined surface was studied by combining the dynamic morphology with the random topography based on empirical data. The results show that the random morphology has a great influence on the flow loss and therefore cannot be ignored. Zhang et al. [18] proposed a milling topography simulation model considering tool wear. The experimental results of cutting under plane and cylindrical surfaces are in good agreement with the model prediction, which proves the correctness of the model. Irene et al. [19] established a numerical model to predict the surface topography of milling process based on the contact relationship between tool and workpiece.

Gao et al. presented an analytical model that considered the effects of tool geometry, cutting conditions, and plastic flow measurements to predict surface topography [20]. Arizmendi et al. established a surface topography prediction model considering tool vibration [21]. Lazoglu et al. [22] established an analysis model for the surface topography of five-axis milling process considering cutting parameters, number of cutting edges, and cutter runout.

The cutting trajectory is another factor that should not be neglected in the prediction model of surface topography. Based on the tool machining paths and the trajectory equation of the cutting edge relative to the workpiece, Zhang et al. [23] developed an iterative algorithm for numerical simulation of machining surface topography in multiaxis milling. Gao et al. [24] proposed a new method to predict the machined surface topography with relative cutting edge trajectory equation. Tan et al. explored the effects of different tool paths on the surface morphology, and the results are shown in Figure 3. By comparing Sa and Rt values, it found that vertical upward is the optimal cutting path [25]. Chen et al. [26] studied the effect of different angle combinations on the surface topography in multiaxis milling. Better surface roughness could be achieved when rotation angles are 0 (positive lead), 60 (combination of positive tilt and positive lead), 90 (positive tilt), and 330 (combination of negative tilt and positive lead). Lu et al. [27] studied the micrometer milling mechanism of Inconel 718 and concluded that the status of the machined surface is determined by the profile of the cutter, the cutting path of the cutter and the flexible deformation of the cutter. Therefore, the author comprehensively considered these three factors to build three-dimensional surface topography and a surface roughness prediction model. Compared with the experimental results, the maximum relative error of the model is 10.9% and the average relative error is 6.8%. The results can be used as a reference for the prediction of surface milling.

**Figure 3.** Effect of different cutter path orientations on surface topography [25]: (**a**) Vertical upward; (**b**) vertical downward; (**c**) horizontal upward; and (**d**) horizontal vertical downward.

By combining boundary intersection and mean squared error method, Brito developed the prediction on surface roughness in AISI 1045 steel end milling process. The results showed that the achieved optimum lessens the sensitivity to the variability transmitted [28]. Tangjitsitcharoen et al. used the dynamic milling force ratio to predict the surface roughness of the ball milling process. The model was verified by experiments. The results show that the model can predict the average surface roughness accuracy up to 92.82% [29]. Also, some other research works were developed to get reasonable surface texture [30,31]. Toh et al. analyzed the surface texture generation mechanism under different tool paths by using the method of surface topography analysis, and obtained the best surface texture when the workpiece was inclined at 75◦ for high speed milling and then determined the optimized machining path [32].

Meanwhile, there may be defect appearing on the machined surface in milling process. Common surface defects are feed mark, tearing surface, and burr formation. Feed marks are produced by the combined effect of tool rotation caused by cutting speed and tool movement caused by feed speed [33]. Feed marks on machined surfaces become more pronounced with increasing cutting speeds and tool wear [34]. Tearing surfaces usually include smeared material, surface microvoids, scratches, and

groove and pitting corrosion. The typical tearing surfaces is showed in Figure 4. Damage on the machined surface in hard milling process of FGH95 PM superalloy was investigated by Du and Liu [35]. Their research results showed that several defects appeared on the machined surface at higher cutting speeds. The results are significant for the prediction of component service life.

**Figure 4.** Three kinds of tearing surface [34]. (**a**) Microvoids and scratches; (**b**) microvoids and scratches; and (**c**) groove and pitting corrosion.

When the tool moves along the feed direction, the workpiece material will flow laterally under the combined extrusion action of the minor flank face and the machined surface causing smearing. The surface microvoids are formed from the carbide particles of the workpiece material, the cutting off of the cutting tools, and the deposition of the chip. The surface microvoids affect the mechanical properties of the workpiece, so the surface microvoids should be avoided in the key components [36].

Also, burr defects are often observed in the area where the tool exits the cutting zone, and the formation of burr is easy to cause stress concentration and reduce the fatigue life of the components. Generally, the materials with low thermal conductivity and Young's modulus are more likely to form burrs on the cutting surface [37]. Under normal circumstances, the burr decreases with the increase of cutting speed, and increases with the increase of tool wear [34].

Generally, cutting parameters with high cutting speed, low feed rate, and low cutting depth is recommend for a good surface finish for milling process.

#### *2.2. Fatigue Strength and Wear Resistance as Effected by Surface Topography*

Milling process is generally applied to conduct finishing machining of sculptured surfaces, therefore the machined component have very demanding specifications in surface topography for its performance. The smoothest surfaces are desired in most milling processes, especially when the fatigue life of the part being machined is high [38]. However, in some biomedical fields, it is desirable to have rough surface morphology [39]. The surface structure is responsible for the mechanical functionality of the component. Sometimes, even if the surface dimension and surface finish of a component are well within the tolerance, there remains the possibility of lack of surface quality for a milled surface. The reason is surface topography influences not only the mechanical and physical properties of contacting parts, but also optical and coating properties of some no contacting components.

Because of the stress concentration caused by pits and groove, the characteristics of milling morphology have important influence on its performance. Generally, rougher surfaces are expected to encourage fatigue crack initiation. It is suggested that parameters such as Rt and Rz are more appropriate than Ra in respect of fatigue strength, as they equate with adverse component surface features [38]. Arola et al. [40] used surface roughness to estimate the effective stress concentration coefficient of high strength low alloy steel workpiece. It was found that the fatigue strength decreased with the increase of surface roughness at low stress level. Moussaoui et al. studied the factors that affect the fatigue life of titanium alloys. They found that the surface roughness does not affect the fatigue life of the workpiece [41]. Guo et al. found that the surface roughness has a slightly influence on four-point bending fatigue life for end milling process of AISI H13 [42]. Li et al. [43] found that surface roughness has little influence on fatigue life due to the small ratio between the ten-point surface height and the bottom curvature radius of surface gaps. Wang et al. [44] found that high surface roughness leads to a high stress concentration coefficient, which reduces the fatigue life of the workpiece.

The machined surface morphology affects the fatigue performance of the final assembly, especially when the crack initiation life is noteworthy [45,46]. Therefore, many works have been done to explore the effect of surface topography on part performance. The traditional two-dimensional surface roughness does not adequately characterize the effect of surface properties on the fatigue performance, while the three-dimensional topography can provide a more accurate correlation with fatigue behavior [47–49].

Performance and reliability of engineering components can be enhanced by selecting appropriate 3D topographic characteristics [50]. Novovicetc et al. investigated the effected of surface and subsurface condition on the fatigue life of Ti alloy workpiece. Their research results indicated that the surface topography, in particular texture direction, showed a strong correlation with the fatigue life of workpiece [51]. Piska et al. [52] studied the effect of progressive milling process on the surface morphology and fatigue properties of 7475-T7351 high-strength aluminum alloy. The results show that the surface topography is not the right factor affecting the fatigue properties of aluminum alloys.

When the roughness level is the same, the milling surface has a higher fatigue limit than the grinding surface. The reason is that the surface grooves produced by milling process are more random [53]. Huang et al. [54] studied the effect of different tool paths on the fatigue life of AISI H13 high speed milling process. It was noted that different tool paths lead to the difference of microscopic stress concentration caused by the orientation morphology, which affects the fatigue performance of the workpiece.

The surface integrity after machining directly affects the life and reliability of the workpiece, so it is important to study the effect of different surface integrity on the friction and wear properties of the workpiece surface to improve the service life of the workpiece. Sedlaek et al. [55] studied the relationship between surface roughness parameters and friction and wear properties of mold surface. The results show that the friction coefficient is inversely proportional to the surface roughness under the condition of dry friction, and the change trend is opposite in the lubrication conditions. Menezes et al. [56] studied the effect of surface texture on friction coefficient and adhesion wear in friction pair. The results show that the friction coefficient and adhesion wear are mainly influenced by the surface texture, which is independent of the surface roughness. Magri et al. [57] studied the relationship between surface morphology and wear resistance during die milling. They found that the best tribological performance was that composed of microcavities generated by similar and high values of fz and ae. The surface topography under four cutting conditions and the die wear results is shown in Figure 5.

**Figure 5.** Four types of surfaces topography and die flash land after forging [57]. Profiles of the die flash land surfaces under (**a**) condition 1, (**b**) condition 2, (**c**) condition 3, and (**d**) condition 4; Die flash land after forging under (**e**) condition 1, (**f**) condition 2, (**g**) condition 3, and (**h**) condition 4.

After milling processes, surface topography has a direct impact on part functional performance, especially in respect of fatigue life and wear resistance. The reason is the surface topography has close relationship with surface frictional characteristic and stress distribution of milled surface. To get a good performance part, the surface topography should be selected according to the application condition of workpiece.

#### **3. Part Functionality as Effected by Surface Metallurgy**

#### *3.1. Machined Surface Metallurgy Characteristic*

In milling of high hardness material or difficult–to-machine metal material, high stress, strain rate, and temperature will have a severe impact on the machined surface. The microscale and nanometer scale existing on the machined surface will change under the interaction of large strain, high strain rate and high temperature [58,59]. The property changes include microstructure change, plastic deformation. Many research works have been done on process mechanics and surface integrity due to the complex coupling between phase transformations and loading in milling process [60]. Generally, the depth of microstructural alteration has been observed to increase when the cutting speed, feed rate and tool wear are increased.

For milling process, Elbestawi et al. investigated the microstructural alterations in high speed milling of hardened AISI H13 using PCBN (Polycrystalline Cubic Boron Nitride) ball-nose end tools. They found that the formation of phase transformation layer was dependent on tool edge preparation and tool wear [61].

Both thermal and mechanical effects attribute to the plastic deformations significantly. Mechanical effects play a major role in the hardening of materials, while thermal effects play a major role in the softening of materials. Moreover, the rubbing effect of flank face on machined surface play an important role in its generation. That is why the tool wear increases from initial condition to its life value, the changed microstructure turns to be deeper. The effects of cutting speed and tool wear on the surface microstructure were investigated when milling titanium alloy Ti-1023 [34]. When VB = 0 (Tool wear value), the cutting speed has no effect on the phase transition or deformation of the machined surface, as shown in Figure 6a. Tool Wear has a significant effect on the plastic deformation and the depth of microstructure of the machined surface, as shown in Figure 6b.

The reasonable selection of cutting parameters in the finishing process does not cause the change of the structure type of the machined surface. The reason is the contact area between the cutting tool and the machined surface is so smaller that the maximal temperature of the machined surface would be lower than the austenitizing temperature of workpiece material [62]. The similar results were found by Li and Zhao [63] and Devillez et al. [64].

Extensive experimental work has been conducted to get an ideal surface metallurgical for machining process. However, indispensable hardware is needed for this method. So the modeling for machined surface microstructure has attracted wide attention [65–69]. Unfortunately, few research works focused on the milling process are found.

After high-speed milling of Ti-6Al-4V and Ti-834, the microstructural subsurface damage in the form of intense slip bands was identified by Thomas et al. Due to a reduction in fatigue crack initiation resistance, the microstructural subsurface damage could degrade the in-service properties of workpiece [70]. Shyha et al. [71] studied the influence law of cutting fluid supply system on metallurgical characteristics, and concluded that cutting fluid had little influence on the microstructure and deformation layer of cutting subsurface; cutting speed was a key factor affecting microstructure. Li et al. [72] carried out an experimental study on hard milling of AISI H13 steel, and the results showed that the nanohardness and plastic deformation depth of the machined surface increased with the increase of the grinding radius of the cutting edge. Liu et al. [73] conducted an experimental study on the AA7150-T651 aluminum alloy. It was found that the severe shear strain caused by the mutual friction between the workpiece and the cutter resulted in a high deformation layer near the surface area.

**Figure 6.** Effects of cutting speed and tool wear on subsurface microstructure [34]. (**a**) VB = 0 m and (**b**) Vc = 60 m min<sup>−</sup>1, fz = 0.08 mm/tooth, ap = 1 mm.

Generally, if the temperature generating on the milled surface is higher than austenitizing temperature of workpiece material, then there will be phase transformations layer appearing on the milled surface. Combining with mechanic load in milling process, the thermal load will make the surface and subsurface metallurgy change. Additionally, the effect of tool should not be neglected. Because the cutting tool will cut-out and the tool will cut-in during the milling process, the thermal and mechanic loads applied to the workpiece do not remain the same. So, the relationship between surface metallurgy characteristic and cutting parameters should be investigated future.

#### *3.2. Wear Resistance and Corrosion Resistance as Effected by Surface Metallurgy*

The microstructural alterations in the material cause the surface layer to exhibit different material behavior, and these alterations include phase transformations and plastic deformations. After the workpiece is machined, the behavior and property of surface is different with the interior of the bulk material. So, the subsurface microstructure has a crucial impact on the performance of the final part. There are controversial standpoints about whether the formation of phase transformations is beneficial to the application. It is more brittle than bulk material, so the appearance of the white layer usually worsens the product's service life [74]; therefore it is important to prevent its occurrence or at least predict how it would affect the final product.

There is similar pattern of higher hardness on the milled surface because of microstructural alteration. In the production process, the effect of hardening layer can be eliminated when the cutting depth is greater than the hardening layer of the workpiece, but it is difficult to realize [75]. Minimizing or eliminating the phase transformations layer would improve the machined surface quality and fatigue strength. Microhardness is a comprehensive index characterizing the microstructure of surface materials, which can be used to characterize the effect of microstructure on fatigue performance. Related research shows that the increase of microhardness in a certain range can improve the fatigue life of workpieces [76,77].

Fellah et al. [78] found that crystallite and grain size play a controlling role in friction coefficient and wear rate. The smaller the grain size, the higher the wear resistance. Zhao et al. [79] studied the friction characteristics of the machined surface of titanium alloy and found that the wear resistance of the workpiece material increases gradually with the increase of hardening degree and grain refinement degree of the machined surface. Huang et al. [80] studied the friction and wear behavior of milling AISI D2 steel. It is found that the subsurface grain deformation induced by machining is helpful to improve the wear resistance of the workpiece. At the same time, it is pointed out that the depth of deformation zone and grain boundary inclination angle can be used as the evaluation index of wear resistance to some extent.

The corrosion resistance of 7050-T7451 aluminum alloy processed by high speed milling was studied. It is found that the surface corrosion damage of the workpiece is determined by the interface energy between the grains and the degree of hardening [81].

Additionally, Kim et.al. used two different methods to estimate the effect of cooling methods on hot forging die service life against plastic deformation and abrasive wear. They found that the die service life depended on abrasive wear, rather than the plastic deformation of the die, for a specific cooling method [82]. To enhance the functional properties of the machined surface, roller burnishing is an effective approach as it changes the microstructure of surface and subsurface [6].

The alterations of surface metallurgy have an important impact on part functional performance. Also the effects of surface metallurgy on part performance are complicated in different application conditions. To get better performance of milled workpiece, the effect of phase transformations and plastic deformations in surface metallurgy should be considered comprehensively.

#### **4. Part Functionality as Effected by Surface Residual Stress**

#### *4.1. Residual Stress Characteristic on Machined Surface*

Residual stresses are stresses that remain in a solid body after the external loading (mechanical and thermal) has been eliminated [83]. At the same time, the microstructure changes of machined surface can also cause residual stress [84]. Some scholars have studied the basic principle of residual stress caused by machining [33,85,86]. As far as 1982, Brinksmeier et al. have conducted such research. They reported the causes of residual stresses in machining process [87]. The final state of residual stress on the component has close relationships with other surface integrity factors, such as surface topography, subsurface microstructure, and topological states of a machined surface.

Considerable research on residual stress has been done for milling process. Titanium alloy is widely used in aerospace industry because of its high mechanical strength, chemical resistance, and thermal conductivity, so it is very important to control the distribution of residual stress. Sun et al. studied the surface integrity in the process of Ti-6Al-4V milling by the experimental method. It was found that the compressive residual normal stresses in the cutting and feed direction increase with the increase of cutting speed. Meanwhile, it is found that the compressive residual normal stress in feed direction is larger than that in cutting direction [88]. Rao et al. studied the milling of Ti-6-Al-4V titanium alloy and concluded that the compressive residual stress increased with the increase of feed and the cutting speed had no effect on the compressive residual stress [89]. For milling process of Inconel 718 with carbide K20, the increase of milling speed will enhance the tensile stress at the surface and compressive stress beneath the machined surface.

Aspinwall et al. used experimental methods to study the effect of tool positioning and workpiece angle on machining performance during milling Inconel 718. Compressive residual stress is generated in horizontal upward cutting and tensile residual stress is generated in horizontal downward cutting. The reason is that horizontal upward has a relatively low cutting speed, so the local temperature of the machined surface is lower. The influence of tool wear and cutting direction on residual stress is shown in Figure 7 [90]. Jiang et al. investigated the effects of tool diameters on the residual stress in the milling process of a thin-walled part. They found that the distribution of residual stress was more uniform as the tool diameter increased [91]. For milling process of nickel alloy and titanium alloy, many research results indicated that increasing cutting speed will bring about the tensile residual stresses tend to

become more compressive [92]. Axinte et al. studied the residual stresses in the feed direction during milling AISI H13. It is found that cutting speed and feed per tooth are important factors affecting the residual stress on the surface [93]. When AISI H13 steel is hard milled by coated cutting tools, the investigated result of in-depth residual stress distribution showed that microstructural changes deeply affect the residual stress and that they have to be accurately taken into account during the process design [94]. Wang et al. [95] studied the effects of cutting parameters, cutting fluid, and spindle angle on residual stress during milling of Inconel 718 alloy. The results show that the cutting depth and cutting speed have great influence on the distribution of residual stress, and the residual stress in the tensile direction increases gradually with the increase of the spindle angle.

**Figure 7.** Effect of different cutter orientations/workpiece tilt angle on residual stress [90].

Although the experimental method can directly reflect the relationship between the distribution of residual stress and cutting conditions, it has higher cost and longer period. Therefore, it is urgent to develop a predictive model of residual stress distribution in milling process. Jiang et al. [96] used finite element simulation to study the distribution of residual stress in high-speed peripheral milling, and found that the cutting thickness had a significant impact on the tangential residual stress. Li et al. [97] used the finite element method to study the effect of cutting depth on the residual stress in milling, and pointed out that the residual stress can be reduced by optimizing the cutting depth of thin wall parts. Liang et al. [98] established an exponential decay function considering the flank wear, tool inclination, and depth of cut to predict the compressive residual stresses distribution of the milled TC17 alloy (An alpha-beta titanium alloy). Fergani et al. [99] used Neumann–Duhamel principle to establish the regeneration prediction model of residual stress in multipass machining under thermomechanical loading. Considering the three-dimensional instantaneous contact state between the milling cutter and the workpiece, Wan et al. [100] established the theoretical model for predicting the residual stress in the milling process. At the same time, it was observed that the thermal load had a relatively weak effect on the residual stress.

Residual stress is a crucial factor to evaluate surface integrity of the milled workpiece. In milling process, distribution of residual stresses is marginally affected by the cutting speed and tool wear state. Also, its residual distribution has close relationship with cutting strategy in milling process. To get reasonable distribution of residual stresses, the cutting parameters should be optimized with consideration of tool parameters.

#### *4.2. Fatigue Resistance as Effected by Surface Residual Stress*

The residual stress has an important effect on the mechanical properties of the workpiece. Residual stresses directly influence the deformation of workpiece, such as static and dynamic strength and chimerical and electrical properties, as shown in Figure 8. It is well known that surface finish and residual stress can significantly affect the antidestructive ability of the component under high cyclic fatigue load.

**Figure 8.** Effect of residual stress on workpiece performance.

Compressive residual stress is usually advantageous to the fatigue life of machined parts, while residual tensile stress is the opposite. The tensile residual stress will enlarge or contribute the extension of microcrack. When the crack increases to a certain extent, it will cause workpiece failure. So, it is necessary to remove tensile residual stresses occurring during machining process. In view of the high speed milling process of Ti-10V-2Fe-3Al, Yao et al. [101] found that the fatigue life of the workpiece is more sensitive to residual stress than surface roughness. Similar results were found by Moussaoui et al. [41].

Influence of milling process on the fatigue life of Ti6-Al-4V was investigated by Moussaoui et al. They found that residual stress has a more preponderant influence on fatigue life than geometric and metallurgical parameters [102]. Huang et al. [55] studied the effect of different tool paths on the fatigue life of AISI H13 high-speed milling process. The results show that the different tool paths lead to the difference of effective residual stress, which results in different fatigue performance of the workpiece. Meanwhile, it is found that the influence of effective residual stress on the fatigue life of the workpiece is greater than that of the orientational surface topography. Yang et al. [103] studied the effect of residual stress on the fatigue life of workpieces during milling of Ti-6Al-4V. The results show that increasing the compressive residual stress can effectively improve the fatigue performance of the workpiece. However, when the workpiece surface material produces plastic deformation, the effect of residual stress will disappear.

Residual stress is a crucial factor to evaluate surface integrity of the milled workpiece. In milling process, distribution of residual stresses is marginally affected by the cutting speed and tool wear state. Also, its residual distribution has close relationship with cutting strategy in milling process. To get reasonable distribution of residual stresses, the cutting parameters should be optimized with consideration of tool parameters. If the distribution of residual stresses is not reasonable, it can be modified due by mechanical loads (static or cyclic) or thermal exposure [104].

The influence of residual stress and microstructure on the fatigue life of the workpiece was studied by F. Ghanem et al. By comparing electro-discharge machining (EDM) with the milling process, they found that the surfaces prepared by EDM showed a tensile residual stress at the surface. The milled surfaces showed a near-surface compressive residual stress, which is favorable for fatigue crack resistance, and the comparison result is shown in Figure 9 [105]. Three-point bending fatigue tests of the notched specimens revealed a loss of 35% in fatigue endurance in the case of EDM.

Tool wear is an unavoidable and complicated phenomenon occurring in machining process, which has a direct impact on residual stress on milled surface. For milling process, many workpiece materials contain carbide particles in their structure. As the cutting tool wears, some carbide particles in workpiece are removed from the machined surface sometimes, then there will be direct effect on the surface quality of the machined surface. Generally, tool flank wear was found to have the most effect on distribution of residual stress in high speed milling process.

**Figure 9.** Comparison effects of residual stress on fatigue life between electro-discharge machining and milling [105].

Considerable research has been performed to get the effect of tool wear on part functionality. The effects of tool wear on surface integrity and fatigue life during milling was studied by Guo et al. They found that surface roughness is generally higher in the step-over direction than the feed direction and that tool wear does not necessarily affect the fatigue life within a certain range (0.2 mm) [106]. Also, fundamental relationship between tool wear and fatigue was related by them. They found that compared with the sharp tool, the worn tool up to flank wear (0.12 mm) only slightly reduces the average life from 1.23 × 106 to 1.16 × 106 cycles. Rougher surfaces are expected to encourage fatigue crack initiation. While Guo found that the surface roughness also has a slightly influence on four-point bending fatigue life for end milling of AISI H13 [42], and the effect of tool wear on workpiece fatigue life is shown in Figure 10.

**Figure 10.** Fatigue life vs. tool wear and fatigue fracture pattern [42].

The residual stress on milled surface will directly influence the generation and extension of crack on surface, and that will have a direct impact on fatigue life of workpiece. Generally, compressive residual stress contributes to the improvement of fatigue life. Tool wear has a direct impact on the residual stress of the milled surface, so tool wear is another important factor that influences the fatigue life of component in milling process.

#### **5. Conclusions and Outlook**

The optimization and control of machined surface integrity is the key technology to ensure the functional performance and service life of parts. The effects of cutting parameters such as cutting speed, feed speed, cutting depth, and tool wear on the machined surface integrity during milling are emphatically reviewed. At the same time, the research progress of the relationship between the machined surface integrity and the component use performance is also revealed. Researchers have made extensive research on the influence of machined surface integrity on fatigue properties and obtained a series of research results. However, there are still some urgent problems to be solved.


**Author Contributions:** The idea of this project was conceived by C.Y. C.Y and H.G. consulted to relevant high-level papers on aspects of changes in surface integrity caused by changes in the milling process, and wrote this project. X.L. and S.Y.L. reviewed this project and proposed constructive guidance to make the article more complete.

**Funding:** This research was funded by the National Natural Science Foundation of China (Grant No. 51575147) and the Science Funds for the Young Innovative Talents of HUST (No. 201507), and the International Cooperation and Exchanges NSFC (No.51720105009).

**Conflicts of Interest:** The authors declare no conflicts of interest.

#### **References**


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