Surface & Coatings Technology 239 (2014) 160–170

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Surface & Coatings Technology

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Surface modification of Ti–6Al–4V alloy via friction-stir processing:Microstructure evolution and dry sliding wear performance

Bo Li a,⁎, Yifu Shen a, Weiye Hu b, Lei Luo a

a College of Materials Science and Technology, Nanjing University of Aeronautics and Astronautics, 29 Jiangjun Road, Nanjing 210016, PR Chinab Technology Research Institute of Nanjing Chenguang Group Co., Ltd (NCGC), China Aerospace Science and Industry Corporation (CASTC), 1 Zhengxue Road, Nanjing 210006, PR China

⁎ Corresponding author. Tel.: +86 15050597986; fax:E-mail addresses: libo_2009@nuaa.edu.cn, libo_2009_

0257-8972/$ – see front matter © 2013 Elsevier B.V. All rihttp://dx.doi.org/10.1016/j.surfcoat.2013.11.035

a b s t r a c t

a r t i c l e i n f o

Article history:Received 15 October 2013Accepted in revised form 18 November 2013Available online 23 November 2013

Keywords:Titanium alloyFriction-stir processingβ-regionsMartensiteWear performanceSurface modification

The friction-stir processing (FSP) method was employed to the surface modification of Ti–6Al–4V alloy applyingdifferent processing parameters. The defect-free friction-stir processed surface layers with ~2.5 mm thickness,with final microstructures of β-regions with acicular-α and GB-α, or Martensite-α′ phase, were obtainable.The FSP processing thermal cycles, macro/micro-structures, phase evolutions, microhardness measurements,and dry-sliding wear performance of the processed surface layers were investigated. Considering to the basiclaws of Ti–6Al–4V microstructure evolutions during the FSP procedures, the relations between processing pa-rameters and produced microstructure characteristics were further elucidated. It was found that a relativelylower processing peak temperature and a fast enough cooling rate benefited the formation of the fine β-regionswith ultra-refined acicular-α phase and/or even needle-typed Martensite-α′. The microstructures favored theimprovement of hardness andwear performance of the processed surface layers, in comparisonwith those prop-erties of the base material.

© 2013 Elsevier B.V. All rights reserved.

1. Introduction

Ti–6Al–4V alloy is the most frequently and successfully used α + βtitanium alloy in various industries due to itsmany favorable properties,including its high strength to weight ratio, low density and biocompat-ibility [1]. Nevertheless, Ti–6Al–4V alloy has poor surface wear proper-ties when used in some harsh environments [2,3]. It limits the productservice life inmechanical friction engineering and transmission compo-nents for human body, such as hip implants [4]. Formany applications itis sufficient or even desirable to reinforce only a surface layer while theother part of the component retains its original composition and struc-ture with high toughness. Considering that the wear performance de-pends on the surface layer, rather than the bulk alloy, the surfacemodification could provide a solution to allow the longevity enhance-ment of Ti–6Al–4V alloy products [5].

Friction-stir welding and processing (FSW/P), as an emerging solid-state joining and manufacturing technique, has already attracted con-siderable attention in multiple industries due to its unique process ad-vantages and high success for joining/processing of many Al and Mgbased light alloys [6–8]. The concept of friction-stir processing (FSP)was first proposed by Mishra et al. [9] and initially employed for Alalloys according to the fundamentals of friction-stir welding (FSW),which is invented by The Welding Institute (TWI) of UK [10]. After re-cent years of research, it has been reported that the coupled thermal–mechanical effect of FSP procedure can be utilized to provide localized

+86 2552112626.nuaa@163.com (B. Li).

ghts reserved.

microstructure modifications in the surface layer of various metals ofAl, Mg and Cu based component, including grain-refining and even pre-paring nano-crystalline [7,11–13]. FSP is performed completely in solid-state, by plunging a rotating, cylindrical toolwith an extended pin into awork piece and traversing it. Therefore, the pin length as a process var-iable can alter the thickness of surface layer processed on themetal sub-strate. The prepared surface layer with new microstructure feature byFSP is very considerable in thickness so that it even can be up to centi-meters. More importantly, other than themelting–solidification surfacetreatment methods, such as laser cladding, the solid-state FSP proce-dure can prevent some metallurgical defects including cracking, poros-ity, anisotropic and dendritic grain coarsening, and excess residualstress after the fusion–re-solidification procedures. In addition, the pro-ductive efficiency of FSP is very considerable on the premise of low cost.Thus, the aforementioned technical advantages of FSP stronglymotivat-ed the research interests on metallic surface modification via thismethod.

FSP has been applied extensively to Al and Mg alloys. However re-cently, with the great progresses in FSW/P tool materials, there hasbeen increasing interest in FSP on the surface of high melting point ma-terials such as titanium, iron and nickel based alloys [14–16]. Due to thatTi has low thermal conductivity and high softening temperature, it isdifficult to generate sufficient heat-input to soften material and realizeconsiderable plasticized deformationwithout causing local overheating.Consequently, understanding of the influences of FSPprocessing param-eters on the microstructure evolutions and the resulted surface proper-ties of Ti and its alloys is quite important. Several literatures on FSP ofTi–6Al–4V alloy, especially for refining the coarseness, fully lamellar

161B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

microstructure of investment-cast and hot isostatic pressed Ti–6Al–4Valloys, have been already published [15,17,18]. However, the relationsamong the FSP processing parameters, microstructure evolution lawsand surface tribological properties are still far from being fullyestablished and understood. We believe that there is merit inattempting to improve those. Additionally, due to the grain-refining ef-fect, FSPmay provide a new, low-cost way to enhance the surface tribo-logical performance of Ti–6Al–4V alloy, without introducing othermaterials to matrix or preparing hard coatings.

In this work, FSP was used to process the surface layer of a commer-cial Ti–6Al–4V alloy. The processing thermal histories or heating–cooling cycles, produced macro- and micro-structures, phase evolu-tions, hardness profiles, and dry sliding wear performances of thefriction-stir processed (FS-processed) Ti–6Al–4V alloy were investigat-ed. The relations between processing parameters and microstructurecharacteristics were further discussed. The basic laws of Ti–6Al–4Valloy microstructure evolutions during the FSP procedures were eluci-dated in detail.

2. Experimental details

Three-millimeter-thick rolled plates of Ti–6Al–4V alloy (6.01 wt.%Al, 3.84 wt.% V, 0.3 wt.% Fe, 0.1 wt.% C and bal. Ti), with α + β duplexphases after annealing treatment, were used as substrates in the exper-imental work. The typicalmicrostructure of un-processed substratewasgiven as Fig. 1. FSPprocedureswere performedusing a professional FSWmachine with an FSP tool of WC–13 wt.% Co matrix material. The con-trol mechanism of the FSW machine was adopted as ‘position-control’mode. As exhibited in Fig. 2a, the FSP tool was composed of a cylindricalshoulder with 15 mm in diameter and a pin in the shape of circulartruncated cone. The surface of shoulder was flat in shape. The pin wasshaped as circular truncated cone, andwas tapered from6 mm in diam-eter at the root to 4 mm in diameter at the pin tip. The pin length was2.2 mm to produce an FS-processed surface layer on the substratewith a thickness of slightly more than 2.2 mm.

During the FSP procedures, a specialized gas protection devicewith agas chamber fixed on the work operating platform was utilized to con-tinuously introduce a shielding atmosphere of argon gas (99.9% purity)into the chamber. As illustrated in Fig. 2b, a sliding panel was settledwith the stationary spindle above the rotation spindle holding the FSPtool. The air impermeability of the gas protection device was well en-sured as possible through a good physical design. Hence, the shieldingargon gas was introduced surrounding the rotating tool and upperpart of the FS-processed zones aiming to prevent the high-temperature oxidation of the Ti–6Al–4V alloy.

20μμm

Prior-α

Prior-β

Fig. 1.Microstructure of BM by SEM.

Furthermore, an infrared temperaturemeasurement systemwas ap-plied to monitor the temperature changes of the substrate surface dur-ing FSP on Ti–6Al–4V alloy plates, aiming to investigate the thermalhistories and cooling rates of the FSP procedures. The infrared tempera-ture probe (IRTP) (as sketched in Fig. 2b) was statically aimed to thesuitable position on the center line of the FS processed surface layeralong the tool traveling direction. The distance between the IRTP aimingpoint and the initial FSP tool plunged position was uniformly set as60 mm for every FSP pass. The obtained real-time temperature valuesfrom IRTP were automatically recorded per 1 s to draw temperature–time curves. For eliminating the physical obstruction of the gas protec-tion device chamber against the infrared point, the IRTP was embeddedwith the gas chamber and placed on the side of the FS processed pass. Itshould be noted that, the fluctuations of temperature–time curves dueto the interference of traveling FSP tool were inevitable and partlyremained.

In terms of metallic surface modification, the multi-pass FSP proce-dures using the consistent processing parameters can be employed toobtain a large scale FS processed zone [7,19]. It means that the micro-structure characteristic in the FS processed zone of every FSP pass isequivalent [7,19]. Thus, in the present study, only the single-pass FSPprocedures on Ti–6Al–4V alloy substrates were investigated. The mainFSP processing parameters included tool rotation speed (n, r/min),tool traveling speed (v, mm/min), tool-shoulder plunge depth (d,mm) and tool tilt angle (α, °). Other than the FSP for Al or Mg alloys,in the present experiments, the d value of FSP tool for Ti–6Al–4V alloyshould be lower as no more than 0.05 mm to reduce the tool-shoulderwear and ensure the required friction heat-input. The α was 0° for awell performed surface in macro scale. After a series of previous pro-cessing experimental works for ensuring a defect-free formability ofFS processed surface layer, the n was tailored as 150–650 r/min, andthe v was tailored as 60–210 mm/min. In this paper, the variations ofFSP processing parameters selected for investigating their influenceson the resultant microstructures and properties were all within thevalue ranges of the aforementioned n and v.

Samples for metallographic examinations were cut, ground,polished, and etched for 10 s with a solution consisting of HF (6 mL),HNO3 (8 mL), and distilled water (86 mL). Macrostructures and low-magnification microstructures of FS processed samples were observedby an optical microscope (OM). The microstructure observations weremainly conducted by a Quanta 200 scanning electron microscope(SEM) (FEI Company, The Netherlands) at 20 kV. Phase identificationwas performed by an X-ray diffractometer (XRD) with Cu Kα radiationat 40 kV and 40 mA, at a scan speed of 4°/min. The Vickers micro-hardness of samples was measured by the HVS-1000 Testing System(Nanjing HVS Inc., China) at 0.2 kg load and 15 s indentation duration.Dry sliding wear testing was conducted in an HT-500 ball-on-disktribometer (LanZhou ZhongKe KaiHua Sci. & Technol. Co., Ltd., PRChina) in air at room temperature. The surfaces of specimens wereground and polished prior to wear tests. A ~3 mm diameter bearingsteel GCr15 ball with a mean hardness of HRC60 was taken as the fric-tion counterface material and a test load of 330 g was applied. The fric-tion unitwas rotated at a speed of 460 rpm for 15 min,with the rotationradius fixed at 2 mm. The coefficient of friction (COF) of the specimenswas recorded during wear testing. The wear volume (V) was deter-mined gravimetrically using

V ¼ Mloss=ρ

where Mloss was the weight loss of Ti–6Al–4V alloy specimens afterwear tests. The density ρ of Ti–6Al–4V alloy was 4.51 g/cm3. The wearrate (ω) was calculated by

ω ¼ V= WLð Þ

r

a

FSPTool

Stationary axle

Gas-out

Gas-in Ti-6Al-4V substrate

Sliding cover

Work table

b

5mm

Shoulder

Pin

Argon atmosphere

IRTP

Fig. 2. View of FSP tool (a), and diagrammatic sketch of the applied gas protective device with a gas chamber with an IRTP (b). Note: the IRTP is an infrared temperature probe.

162 B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

where W was the friction contact load. L was the sliding distance afterthe 15 min duration.

3. Results and discussion

3.1. Surface macroscopic formability and cross-section macrostructure

Before a series of processing experiments for obtaining a defect-freeFS processed surface layer, the process principles of FSP on Ti–6Al–4Valloy should be understood. First, the rotating tool was slowly insertedinto the substrate on the center line. Thus, the localized heats were gen-erated due to the friction between the interface of tool and substrate.After the rotating tool-shoulder fully contacted with the substrate sur-face, the tool traveled along the center line at a certain speed. Duringthe rotating and traveling of tool, the tool behavior of friction and stir-ring was maintained to generate heat-input into the FS processed

c

d e

(d)SNZ

BM

AS

SNZ

BM

TZ

a

Key-holeSurface c

Ring-like textures

TZ

100μm

Fig. 3. Top view of the FS processed surface layer with surface crowns (a) and without surfacelayer (c) and microstructures of TZ at AS (d) and RS (e) by OM. Note: the explanations of abbr

zone and induce the severe plasticization of the Ti–6Al–4V alloy. Infact, during the tool traveling, the main heat resource included notonly the friction heating but also the plastic-deformed energy releaseof the matrix itself [6]. Thus, a long strip of FS processed zone was pro-duced on the substrate via a single-pass FSP procedure after the transla-tional motion of the rotating tool. For the metallic surface modification,the FS processed zonewas then theproduced surface layer,with a thick-ness of slightly more than pin length. A key-hole was remained at thefinal position of FSP pass after the elevating of tool-pin (as marked inFig. 3a).

Fig. 3a exhibits the macroscopic top view of the single-pass FS proc-essed surface layer, using the n value of 350 r/min and v value of60 mm/min. Numerous ring-like textures and unilateral surface crownswere presented. When the vwas increased, or the nwas decreased, thesurface crowns disappeared, as shown in Fig. 3b. It was also one of re-sults from process optimizations. During FSP, the Ti–6Al–4V alloy

1mm

(e) RS

SNZ

BM

TZ

b

5mm

rowns at AS

5mm

TZ

100μm

λ

crowns using a tailored λ value (b); typical cross-section view of the FS processed surfaceeviations in this figure are given in the paper.

163B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

underwent severe plasticized deformation because of the tool rotatingand traveling behaviors at the heat elevating situation. Thus, the soft-ened and then severely plasticized material ahead the traveling toolshoulder was migrated from the front at the advancing side (AS,where the tool rotating directionwas the samewith the traveling direc-tion) to the rear at the retreating side (RS, where the tool rotating direc-tion was opposite against the traveling direction). This migration orflowing behavior of the circle-form, softened band-layer per rotationof tool-shoulder contributed to form the surface ring-like textures. Ac-tually, the two adjacent and overlapping ring-like textures, or semicycleband-layer, on the FS processed surface represented two situations oftool-shoulder at one rotation and then at the subsequent one. Therefore,asmarked in Fig. 3b, thewidth of a ring-like texture per rotation of tool-pinwas defined as λ, which equaled the distance between two adjacentand overlapping ring-like textures. The relationship of λ and FSP pro-cessing parameters was stated as

λ ¼ vn

ð1Þ

where v is the tool traveling speed and n is the tool rotating speed.Aiming to eliminate the unilateral surface crowns (Fig. 3a) and ob-

tain the dense and regular ring-like textures on the FS processed sur-face, the λ value was found as the key factor to be controlled. Thisexperiment work found that when the λ value was suitably in therange of 0.25–0.75 mm, the unilateral surface crowns would disappearon the premise of no macroscopic defect produced in the FS processedzone.

The v, n, and their ratio value of λ were found to significantly affectthe FSP surface quality, which was important to the practical applica-tion. Accompanying every rotation action of FSP-tool, the softened-Tiband-layers were orderly migrated from the tool AS to its RS. The sur-face ring-like textures were composed of the numerous circle-form Tiband-layers closely arranged one by one. The mean width of per ring-like band-layer,λ, was approximately equal to the tool traveled distanceper rotation. In fact, if the n value was invariable, a too low v value re-sulted in the narrowing per band-layer, or even the stacking of severalband-layers at one position. It should be noted that the stacking of sev-eral band-layerswouldfinally form the unilateral surface crowns,whichwas shown in Fig. 3a. An excess v value easily deduced a discontinuousannular morphology on surface, even a surface groove defect. If the vvalue was invariable, an over-slow n value went against the frictionheating generation between tool shoulder and its lower Ti substrate. Itnegatively affected the formability of FSP. An over-elevated n value atan invariable v would also result in the stacking of several band-layersor overheating defects in the processed surface layer. Therefore, an op-timized and suitable ratio value (λ) of v and nwas necessary for a regu-lar, successive and compact ring-like texture on the FS processedsurface. Besides, the surface texture could also be removed by mechan-ical milling or grinding for a stricter requirement for the surfaceroughness.

Fig. 3c shows a typical cross-section macrostructure of the FS proc-essed surface layer. Three zones were presented: stir nugget zone(SNZ), base material (BM) and transition zone (TZ) between SNZ andits adjacent BM. Fig. 3d and e gives the OM images of TZ at AS and RS.It seemed that the width of the band-like TZ at AS was narrower thanthat at RS. The mean grain size in either SNZ or TZ was much smallerthan that in BM. It was the result of the microstructure refining effectof FSP. We believe that the studies on microstructure and propertiesof SNZ were very important, because the bowl-shaped SNZ was almostthe total part of the FS processed surface layer.

3.2. Thermal histories during FSP

It was found that for the most of Al alloys, the maximum processingtemperature of FSW/P generally conformed to a certain relationship

with the process parameters. It can be explained by [6,20]

TTm

¼ Kn2

v� 104

!α

ð2Þ

where the exponent α and the constant K are material parameters, nand v are the rotation and travel speeds during FSW/P, and Tm (°C) isthe melting point of the alloy. The maximum temperature (T) has thepositive correlations with the pseudo-‘heat index’ of n2/v. It suggeststhat the influence of increasing rotation speed on the peak processingtemperature is greater than the decreasing travel speed.

Despite that above relational expression was established on lots ofexperiments for Al alloys, some similar relations between processingparameters and thermal histories of Ti–6Al–4V alloy processed by FSPwere found as well in this work. We believe that the detections andanalyses of thermal cycles including heating–cooling procedures of theprocessed surface benefited investigation on the microstructure evolu-tion laws of the FS processed Ti–6Al–4V alloy. To clear the temperaturemonitoring conditions, Fig. 4a illustrated the placement of IRTP and theFS processed pass. Fig. 4b, c and d gives real-time temperature curvesunder different condition groups of variable FS processing parameters.

As indicated in Fig. 4b, the peak temperature during FSP on the Ti–6Al–4V substrate was generally in direct proportion to the n valuewhen the v value was invariable. The peak processing temperaturesunder the three processing conditions, marked in Fig. 4b, were all obvi-ously higher than the α/β phase transformation temperature (βTransus) of the Ti–6Al–4V alloy. When the n was 650 r/min, the peakprocessing temperature detected on the FS processed surface wasmore than 1300 °C. Due to the invariable v value as 90 mm/min, thecooling rates through the start Martensite transformation temperatureline (Ms temperature) were similar under different n values (as dottedarrows in Fig. 4b). It was found that the cooling rate after the tool trav-eling was strongly related to the tool leaving speed from the FS proc-essed surface behind it. Fig. 4c indicates the real-time temperaturecurves when FSP was performed using different v values and invariablen. Although the peak processing temperatures had no obvious changes,the cooling rate ascended with the increase of v (as dotted arrows inFig. 4c). Considering that the λ value (the FS processed surface ring-like texture width) was strongly related to the formability of a consis-tent surface morphology in macroscopic view by FSP proceduresunder different processing conditions, Fig. 4d indicates the changes incomparisons of not only peak processing temperatures but also coolingrates, when λ was invariable using different values of n and v. Thechanges of both peak processing temperature and cooling rate werethen presented. It implied that the higher peak processing temperatureand higher tool traveling speed favored the elevating of cooling rate.

3.3. Microstructure characteristics under different processing conditions

The BM used in this study was composed of a large number of rela-tively equiaxedα-grains with small amounts of grain-boundary (GB) βphase (Fig. 1). The mean grain size of BM, mainly α-grains, was deter-mined by a linear intercept method at ~22 μm. Fig. 5 exhibits the finalmicrostructures in SNZ produced using different n values but the samev value. Temperature measurements (Fi.g4b) and microstructure char-acterization demonstrated that the α/β phase transformation hadbeen significantly accomplished in SNZ, under these processing condi-tions. The final microstructures contained numerous β grains, or calledβ-regions, with GB-α phase and acicular-α phase emanating from theGB-α into the β-regions (Fig. 5). When n was 350 r/min, the averagegrain size of the β-regions was approximately 15–20 μm (Fig. 5a, b). Ithad been found that the duration above β Transus became longer asthe increase of n value (Fig. 4b). The higher peak temperature duringFSP procedures promoted the growth or coarsening of β grains above

0 20 40 60 80 100

0

200

400

600

800

1000

1200

1400

Tem

per

atu

re (

oC

)T

emp

erat

ure

(oC

)T

emp

erat

ure

(oC

)

Time (sec)

0 20 40 60 80 100Time (sec)

0 20 40 60 80 100Time (sec)

350 r/min; 90mm/min 500 r/min; 90mm/min 650 r/min; 90mm/min

0

200

400

600

800

1000

1200 350 r/min; 90mm/min 350 r/min; 150mm/min 350 r/min; 210mm/min

0

200

400

600

800

1000

1200 150 r/min; 90mm/min 250 r/min; 150mm/min 350 r/min; 210mm/min

transus

FSP tool

Substrate

FSP surface layer

IRTP on side

Travel direction

60mma

(b) Fluctuations due to the interference of traveling tool

c

d

transus

transus

β

β

β

Fig. 4. FSP thermal cycles by temperature measurements: the placement sketch of IRTPand FSP pass (a); thermal cycles using different tool rotation speed values but an invari-able travel speed (b); thermal cycles using different tool travel speed values but an invari-able rotation speed (c); thermal cycles using different values of tool rotation speed andtravel speed but an invariable λ value (d). Note: slope of the dotted arrows representsthe cooling rates through β Transus.

164 B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

the β Transus. Thus, as implied in Fig. 5c and d, the average grain size ofthe β-regions increased using higher n values.

In contrast, as exhibited in Fig. 6, the average grain size of β-regionsin SNZ further but slightly decreased when the v value became consid-erably higher, under the condition that the n value was invariable as350 r/min. That was due to the contraction of time of the high-temperature duration above the β Transus (Fig. 4c). The β-regions

shown in Fig. 6were not perfectly equiaxed and exhibited some elonga-tions. It implied that the higher v played an effect on the resultant exces-sive deformation of β-grains during FSP and the final morphology ofelongated, refined β-regions in SNZ after cooling. Final SNZmicrostruc-ture features of interest included not only β-region grain size but alsomorphology, fraction and distribution of α phase. As shown in Fig. 6aand b, the amount or phase fraction of acicular-α phase within the β-regions obviously became smaller when v increased as 150 mm/minthan that when v was 90 mm/min (Fig. 5a, b). In addition, in compari-son to the α phase morphology shown in Fig. 5, the appeared acicular-α phase within β-regions in Fig. 6a and b was further refined andnarrowed in width. According to the acceleration of cooling ratesusing the higher v values (Fig. 4c), in the quenching-like procedures,α phase had no enough time to be completely transformed. Thus,more β phase was remained with less acicular-α after the cooling.

A more interesting phenomenon occurred in the SNZ when a muchhigher v value of 210 mm/min was adopted using the same n value of350 r/min. As shown in Fig. 6d, the magnification in Fig. 6c, the ultra-refined, needle-type Martensite-α′ phase was formed within the β-regions. It is known that the Martensite-α′ formation in Ti–6Al–4Valloy requires the fairly fast cooling rates through Ms temperature,such as those during water quenching [21]. Esmaily et al. [22] had alsofound that in the friction stir welded Ti–6Al–4V alloy, the Martensite-α′ formation occurred in the areas processed with lower temperaturesbut as a result of faster cooling rates. This microstructure of mixed re-fined α′ and α is extremely small and acicular, which has some specificmechanical properties like high strength and hardness and relativelylow toughness [23]. The Martensitic formation and its differentiationfrom Widmansätten/basketweave structure in the Ti–6Al–4V FSP zoneshould be explained. The valuable information on the diffusional anddiffusionless transformations of Ti–6Al–4Vwas previously documentedby Ahmed and Rack [24]. Ahmed and Rack [24] have shown that inorder to attain different microstructures, the Ti–6Al–4V alloy has to ex-perience different values of cooling rates and this is presented in a formof a continuous cooling transformation (CCT) diagram. According to theCCT diagram for Ti–6Al–4V alloy, it is indicated that the cooling rates inexcess of 410 °C/s through Ms temperature are required to produce afullyMartensitic structure, and that a cooling rate lower than 20 °C/s re-sults entirely in diffusional transformations [24]. A partial isothermaltime–temperature transformation diagram for Ti–6Al–4V alloy hasalso been published, which suggests that a cooling rate in excess of120 °C/s through Ms Transus is required for any Martensite-α′ phaseto form [25]. Although the cooling rate fast enough to form a fully Mar-tensitic structure was not obtained in the present work, the fast coolingrate to form a small amount of Martensite-α′ phase was obtainable onthe processing condition that the n and v were separately 350 r/minand 210 mm/min. On that situation, the cooling rate through Ms tem-perature had significantly exceeded 120 °C/s (1032 °C at 20 s, 874 °Cat 21 s, and 735 °C at 22 s) according to the temperature monitoring(Fig. 4c). However, considering the actual experimental procedureused in the work, especially at high temperatures such as those mea-sured in that setting, the IRTP temperature measurements might be un-reliable. Some special care should be taken to calibrate the instrumentproperly in the same condition as those that did exist in the experiment.Therefore, the cooling rates of the Ti–6Al–4V FSP were compared withthe previously documented CCT diagram, as schematically drawn inFig. 7. A similar drawing about that was also performed in Ref. [22] forthe explanation of the Martensitic formation on the FSW zone of theTi–6Al–4V alloy. Three types of microstructure (i.e. fully Martensitic ordiffusionless structure, diffusional transformations and a combinationof these two) can be formed [22,24]. The generally measured FSPcooling rate in the work, on the processing condition of 350 r/min and210 mm/min, was obviously higher than 20 °C/s but lower than410 °C/s. The approximate value of the FSP cooling rate proved the sug-gestion about the formationofMartensite-α′ phase in the FSPprocessedsurface layer. Thus, the mixed ultra-refined Martensite-α′ and ultra-

-region

-region

GB-α α among β-regions

Acicular- within -regions

Acicular-

GB-

(b)

-region

GB-

Acicular-

a b

c d

-region

20μm 10μm

20μm 20μm

β

β

α β

β

α

α

αα

β

β-regions with acicular-α and GB-α

Fig. 5. SEM images of the FS processed microstructures in SNZ, using the processing parameters of n = 350 r/min, v = 90 mm/min (a and its local magnification of b); n = 500 r/min,v = 90 mm/min (c); and n = 650 r/min, v = 90 mm/min (d).

165B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

refined acicular α phase was produced within and/or among β-regionsin the SNZ of Ti–6Al–4V alloy (Fig. 6d).

Fig. 8 shows the microstructures in the FS processed surface layerusing different processing parameters and the invariable λ value. Bythe microstructure observations, it was found that the mean grain sizeof the β-regions slightly increased as the rise of n value. On the otherhand, α phase within the β-regions significantly became narrower inwidth with the increase of v value. On the premise that the λ valuewas invariable, the suitable rise of cooling rate benefited the formationof ultra-refined acicular-α phase and even the needle-shapedMartensite-α′ phase within the well refined β-regions.

Typical XRD spectra within a wide 2θ range (30°–90°) and a narrow2θ range (39.5°–41°) of the FS processed surface layer of Ti–6Al–4Valloy are depicted in Fig. 9. Diffraction peaks corresponding to hexago-nal close-packed (hcp) Ti and body-centered cubic (bcc) Ti were gener-ally detected in BM and SNZ produced using the two different FSPprocessing parameters (as marked in Fig. 9). Due to that the β phasefraction of BM is very small, there is no obvious bcc Ti diffraction peakmarked in the XRD spectrum of BM (Fig. 9a). However, after FSP, theβ phase fraction became higher according to the intensity changes ofthe appeared bcc Ti diffraction peaks in Fig. 9a. Considering the typicalmicrostructure observation of SNZ produced using n of 350 r/min andv of 210 mm/min, the β phase fraction was indeed higher than that ofBM (Figs. 6, 8, 1). One standard diffraction peak for hcp α-Ti, locatedat 2θ = 40.177° (~40.18°), was taken for a comparison as indicated in

Fig. 9b. At a relatively low v of 90 mm/min, the 2θ location of hcp Ti dif-fraction peaks of the FS processed surface layerwas entirely the same asthat of the standard α phase (Fig. 9b). When the applied v value in-creased to 210 mm/min, it was found that the obtained 2θ location ofthe diffraction peak for hcp Ti generally shifted to a higher 2θ location.Combined with the observations on the SNZ microstructure (Figs. 6d,8c), the shifted hcp Ti diffraction peak corresponded to the status ofMartensite transformation. Additionally, the diffraction peak for hcpα′ phase became broadened. As aforementioned, the acceleration ofthe FSP cooling rate resulted in theMartensite-α′ transformationwithinthe final remaining β-regions. According to Bragg's law, normally, theoccurrence of Martensite transformation is accompanied by microscop-ic volume expansion, which in turn creates stresses at the grain bound-aries that may influence the lattice parameters [26,27]. The observedincrease of 2θ indicated a decrease in theα phase lattice plane distance,which was believed to be mainly caused by the Martensitetransformation.

3.4. Hardness and wear performance

Considering the FS processed microstructure characteristics includ-ing grain refining, α/β transformation, and formation of Martensite-α′, the hardening effect of the Ti–6Al–4V surface layer produced byFSP draws our attention first. As depicted in Fig. 10, the micro-hardnessprofiles on the polished sections of the FS processed surface layer were

(d)

GB-

Refined acicular-

-regions with poor acicular- phase

-region

-region-regionGB-

Martensite-

a b

c d

(b)

β

β

α

α

α

α

α

β

20μm 10μm

20μm 10μm

Fig. 6. SEM images of the FS processedmicrostructures in SNZ, using the processing parameters of n = 350 r/min, v = 150 mm/min (a and its local magnification of b); n = 350 r/min,v = 210 mm/min (c and its local magnification of d).

166 B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

measured, perpendicular to the FSP tool travel direction. The hardnessof BM was 315–320 Hv0.2. That of SNZ was 335–365 Hv0.2 after FSP.When v value was invariable as 90 mm/min, the average hardnessvalue of the SNZ reduced with the increase of n value, which was the

Fig. 7. Schematic continuous cooling diagram for Ti–6Al–4V. The generally measuredcooling rate associatedwith FSP is indicated, proving that theα′Martensite formation oc-curred in the FSP zone at the elevated traveling speed: (i) above 410 °C/s, full-Martensitestructure forms; (ii) amixed combination of non-Martensitic andMartensite structure ap-pears at cooling rates between 410 °C/s and 20 °C/s; (iii) and below 20 °C/s;Widmansätten or basketweave structure is expected. Note: refer to Refs. [22,24].

result of coarsen-grained β-regions in SNZ. Under the processing condi-tions that the n and v were 350 r/min and 150 min/min, respectively,the average hardness in SNZ was slightly lower than that using the nvalue of 350 r/min and v value of 90 mm/min (Fig. 10). It was due tothe significant reduction of the α phase fraction in the β-regions, asshown in Fig. 6a and b. In general, the FSP produced dual α + β phasemicrostructure with fine β-regions and the lamellar α/β characteriza-tion had higher hardness, like that formed under the FSP conditionsthat the n and v were separately 350 r/min and 90 min/min (Fig. 5aand b). Moreover, as exhibited in Fig. 10, the average hardness inthe SNZ processed using the n and v values of 350 r/min and210 min/min, respectively, was slightly higher than the lamellar α/βmicrostructure, due to the significant formation of numerous needle-shaped and ultra-fined Martensite-α′ phase in the further refinedβ-regions (Figs. 6d and 8c). The retention of a reasonable level of re-sidual stress via the Martensite transformation in the processed sur-face layer favored the enhancement of hardness. The significantgrain-refinement effect due to rapid cooling rate and lower process-ing peak temperature also favored a further increase in the hardness.

Fig. 11 depicts the variations of COF with sliding time for the speci-mens of BM, the FS processed surface layer using processing parametersof n and v values of 350 r/min and 90 mm/min, and parameters of n andv values of 350 r/min and 210 mm/min. The wear rates of the threespecimens were also given and recorded in Fig. 11. Detailed SEM obser-vations of the corresponding worn surfaces are shown in Fig. 12. It was

Refined acicular

Coarsening acicular

Martensite-

-regions with acicular- and GB-

a

b

c

α

α

α α

α

β

10μm

Fig. 8. SEM images of the FS processedmicrostructures in SNZ, using the invariableλ valueof 0.6 mm, and processing parameters of n = 150 r/min, v = 90 mm/min (a);n = 250 r/min, v = 150 mm/min (b); and n = 350 r/min, v = 210 mm/min (c).

167B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

found that the BM specimen exhibited a fluctuating state of COF curve,in particular after 9 min sliding duration, and proposed a high COFvalue. However, the FSP procedures exerted significant influences onboth the steady-state behavior and the value reduction of COF curves(Fig. 11). As shown in Fig. 12a, the BM specimen topographical featuresof the worn surface consisted of deep grooves representing abrasionwear. The severe local surface plowingwith some irregular shaped tear-ing pits, which reflected the existence of severe spalling and delamina-tion of the tribolayer, was revealed (Fig. 12a). In addition, the poortribological property of the BM specimen resulted in a high averageCOF value of ~0.56, and a considerably elevated wear rate of 3.96 ×10−4 mm3 N−1 m−1. For the FS processed specimen using n and v

values of 350 r/min and 90 mm/min, the mean COF value was ~0.38with the wear rate of 2.27 × 10−4 mm3 N−1 m−1. As exhibited inFig. 12b, the worn surface contained much narrower grooves, free ofany irregular shaped tearing pits or severe spalling, thereby reducingthe mean COF and attendant wear rate, respectively. Interestingly,using a higher v value of 210 mm/min, theworn surface of FS processedspecimen became smoother with significantly narrow and shallowgrooves, as shown in Fig. 12c. The lowest mean COF of 0.33 and wearrate of 1.25 × 10−4 mm3 N−1 m−1 were accordingly obtained(Fig. 11). In this situation, due to the formed ultra-fined Martensite-α′phase within the well refined β-regions in the FS processed surfacelayer and the attendant enhancement of hardness, the wear propertieswere further improved. Additionally, it was well known that bcc Ti hada better plasticity than hcp Ti. The FS processed microstructures shownin Fig. 6b and d demonstrated that the β phase fraction in SNZ effective-ly increased, while the α phase fraction considerably reduced. There-fore, the plastic protective tribolayer with a large amount of β phasebut with poor acicular-α phase (Fig. 6) easily smeared plastically onthe surface being slid. It was thus reasonable to consider that themech-anism of material removal during sliding changed from abrasion(Fig. 12a) to adhesion of the more plastic tribolayer (Fig. 12b, c). Sucha transition favored the reduction in wear rate after sliding.

3.5. Microstructure evolution laws and the effects on performances

Fig. 13 can be used as aids in the discussion of Ti–6Al–4Vmicrostruc-ture evolution during FSP. Based on the thermal cycle feature, the mapgiven as Fig. 13 can be divided into two parts, heating part and coolingpart, as marked in Fig. 13. The BM contains a large amount of α phasewith little GB-β phase. During FSP procedures, α + β grains are elon-gated and have plasticized deformation due to the high-speed rotatingbehavior of the FSP tool, under the thermal condition provided by thetool/substrate friction heating. In the period of heat elevating, theprior-β phase is further coarsened. Until the period of heat durationabove β Transus, the prior-α phase is transformed to β phase. Afterthe recovery and coarsening procedures, the completely transformedβ grains normally exist when the heating temperature is above βTransus. The growth or coarsening of the transformed β grains can con-tinue until α phase begins to nucleate at β boundaries to limit the GBmigration, during the cooling process below the β Transus. After thegrowth of the transformed α phase in the cooling period, the final β-regions with acicular-α and GB-α will be obtained. If the cooling rateis fast enough, the transformedα phase will have no time to completelygrow; instead, the needle-shaped Martensite-α′ will be produced.

For the FS processed surface layer or SNZ of the Ti–6Al–4V alloy,finalmicrostructure characteristics of our interest include grain size andmorphology as well as the type of phases, phase fractions, and distribu-tion. Based on the analyses of SNZ microstructures, thermal histories,hardness and wear properties under different processing conditionsby FSP, the evolution laws and the effects on the related performanceswere proposed to be as follows, respectively from three aspects of βphase, α phase, and Martensite-α′.

3.5.1. (i) β phaseOther than the small-amount prior GB-β in BM, the β phase fraction

in SNZ via FSP is considerable. The thermal histories obtained throughtemperature measurements prove that the SNZ experiences peak tem-peratures above the β Transus. Moreover, the absence of any retainedprior-α phase in the SNZ indicates that the durations above the βTransus during FSP procedures are sufficient to allow complete transfor-mation to β phase, resulting in the formed final β-regionswith acicular-α andGB-α. As a hotworkingprocess, FSPprocedures exert a significantinfluence on the grain size of the final β phase. The elevated tool rota-tion speed results in a higher peak processing temperature and a longerduration above theβ Transus, which directly promotes the growth of β-grains in the temperature-rise period and the subsequent β-grain

a b

Fig. 9. XRD spectra of BM and the FS processed surface layers within a wide 2θ range of 30°–90° (a), and a narrow 2θ range of 39.5°–41° (b).

168 B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

coarsening in high-temperature duration above the β Transus. Howev-er, the influence of tool travel speed on the β-grain size is not big. Theobtainable well refined β-grains or β-regions of SNZ in the present ex-perimental work are 15–29 mm in size using the tool rotation speedof no more than 350 r/min. Besides, the transformed GB-α during thecooling periods after tool traveling also exerts a limit effect of graincoarsening of β-grains. It is true that the refining of β-regions in SNZbenefits the hardness elevating of the FS processed surface layer. Forthe wear performance, the larger β phase fraction favors the formingof more plastic tribolayer on the surface being slid, to reduce severelocal spalling and surface grooves.

3.5.2. (ii) α phaseDue to the complete α/β transformation in the high-temperature

duration above the β Transus, the acicular-α within β-regions and GB-α that existed in the final SNZ are both formed during the cooling pe-riods below β Transus. Thus, the cooling rate is a key factor of the influ-ence on the grain size, morphology and phase fraction of the final αphase in the SNZ. The GB-α phase is inevitably and first produced be-cause the transformed α phase should preferentially nucleate at triplejunctions and boundaries of β-grains with a low-energy orientation re-lation during the cooling process. The amount of acicular-α phase with-in β-regions can be significantly variable according to the different

310

320

330

340

350

360

370

380

Mic

roh

ard

nes

s (H

v 0.2

)

350r/min 150mm/min 350r/min 210mm/min

-15 -10 -5 0 5 10 15Distance from the center line (mm)

350r/min 90mm/min 500r/min 90mm/min 650r/min 90mm/min

SRtaMBSAtaMB

SNZ

Fig. 10.Hardness measurements of BM and the FS processed surface layers using differentprocessing parameters.

cooling rates. Fast cooling rate results in a reduced α phase fraction. Inaddition, the ultra-fine grain size of α phase is obtainable using thetool travel speed ofmore than 150 mm/min in the presentwork. In con-trast, the lower cooling rate will inevitably generate α grain coarseningand α phase fraction increasing. The produced final α/β lamellar struc-ture in fine grain-size enhances the hardness of the FS processed surfacelayer due to the dual-phase strengthening effect of the α + β Ti alloy.More ultra-refined acicular-α phase within small-sized β-regions is be-lieved to be themain reason of surface hardening. Particularly, althoughthe α phase fraction in SNZ can be significantly reduced by an elevatedcooling rate because the acicular-α has no enough time to be complete-ly transformed, the formation of the Martensite phase in refined β-regions also plays an important role in maintaining a high hardnesslevel of microstructure.

3.5.3. (iii) Martensite-αThe cooling rate through the Ms line during the thermal cycle is

within the limits of 20 °C/s–410 °C/s, using the FSP parameters of350 r/min and 210 mm/min. Research results of the Ti–6Al–4V alloyCCT diagram of Ref. [24] are summarized as follows: cooling rates in ex-cess of 410 °C/s are required for the Ti–6Al–4V alloy to attain acompletely Martensitic microstructure, while lower cooling ratesabove 20 °C/s result inmixedmicrostructures ofMartensiticα′, massive

0 3 6 9 12 150.20

0.25

0.30

0.35

0.40

0.45

0.50

0.55

0.60

0.65

0.70

CO

F

Sliding time (min)

SubstrateFSPed, 350r/min 90mm/minFSPed, 350r/min 210mm/min

Fig. 11. COF with sliding time curves of BM and the FS processed surface layers. Note: thelabeledω1,ω2, andω3 are thewear rate values of Ti–6Al–4VBM, FS processed surface layerusing 350 r/min 90 mm/min, and that using 350 r/min 210 mm/min.

944 m

353 m

292 m

a

b

c

200μm

100μm

100μm

μ

μ

μ

Fig. 12. Wear surface topographies of BM (a); FS processed surface layers usingn = 350 r/min, v = 90 mm/min (b), and n = 350 r/min, v = 210 mm/min (c).

Fig. 13. Diagrammatic sketch of microstructure evolutions of Ti–6Al–4V during FSP.

169B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

α and various morphological forms of diffusion controlled α. Althoughthis cooling condition generally measured by IRTP on the FSP processedsurface layer is not enough for a fully Martensite transformation, theroughly obtained cooling rate provides an important possibility offorming a small amount of Martensite phase in the surface layer of Ti–6Al–4V alloy, according to the SEM, XRD, micro-hardness and dry-

sliding wear analysis results. Under that FSP processing condition, theproduced needle-shaped and considerably ultra-refined Martensitephase favors the improvement in hardness and wear performance ofthe processed surface layer.

4. Conclusions

(1) The defect-free FS processed surface layer of the Ti–6Al–4V alloywas obtainable by the applied processing parameters. A suitableratio value (λ) of the tool travel speed (v) to tool rotation speed(n) benefited the formability of well-distributed ring-like tex-tures on the processed surface pass. Peak processing tempera-tures of FSP procedures exceeded the β Transus. The elevatedtool rotating speed resulted in a higher FSP peak processing tem-perature. The cooling rate through the Ms was accelerated withthe increase of tool travel speed.

(2) The produced final microstructures in SNZ mainly composed ofβ-regions with acicular-α and GB-α. The mean grain size of β-regions reduced under the conditions of lower tool rotationspeed and higher tool traveling speed. The final α phase fractiondecreased when the cooling rate was fast enough. A higher tooltraveling speed favored the further refining of the acicular-αphase within β-regions. Instead of a large amount of acicular-α,ultra-refined and needle-type Martensite-α′ phase was trans-formed in β-regions using the FSP processing parameters of350 r/min and 210 mm/min.

(3) For the FS processed surface layer, the microstructure character-ization of further refined β-regions with ultra-refinedα phase orMartensite-α′ phase was benefited to the improvement of inhardness and dry sliding wear performance, in comparisonwith those properties of BM. The formation of an adherent, plas-tically smeared tribolayer on the worn surface of rich β phase,which was formed in the FS processed surface layer, improvedthe wear property. A group of optimal technical parameters of

170 B. Li et al. / Surface & Coatings Technology 239 (2014) 160–170

350 r/min and 210 mm/min (λ = 0.6 mm) was finally tailoredaccording to the good surface formability, higher mean micro-hardness and satisfactorywear performance of the produced sur-face layer.

Acknowledgements

This work is supported by Funding for Outstanding Doctoral Disser-tation in NUAA (BCXJ12-09), the Fundamental Research Funds for theCentral Universities, PR China.

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