Current research and development status of dissimilar materials laser welding of titanium and its alloys

Current research and development status of dissimilar materials laser welding of titanium and its alloys

Optics and Laser Technology 126 (2020) 106090 Contents lists available at ScienceDirect Optics and Laser Technology journal homepage: www.elsevier.c...

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Optics and Laser Technology 126 (2020) 106090

Contents lists available at ScienceDirect

Optics and Laser Technology journal homepage: www.elsevier.com/locate/optlastec

Review

Current research and development status of dissimilar materials laser welding of titanium and its alloys

T



M.M. Quazia, , M. Ishaka, M.A. Fazalb, A. Arslanc, Saeed Rubaieeb,d, Abdullah Qabane, M.H. Aimana, Tipu Sultanf, M.M. Alig, S.M. Manladanh a

Faculty of Mechanical and Automotive Engineering Technology, Universiti Malaysia Pahang, 26600 Pekan, Pahang, Malaysia Department of Mechanical and Materials Engineering, University of Jeddah, Saudi Arabia c Department of Mechanical Engineering, COMSATS University Islamabad, Sahiwal Campus, Sahiwal 57000, Pakistan d Department of Industrial and Systems Engineering, University of Jeddah, Saudi Arabia e Department of Mechanical Engineering and Aeronautics, City, University of London, London, UK f Department of Restorative Dentistry, Faculty of Dentistry, University of Malaya, Kuala Lumpur, Malaysia g Optical Fibre Sensors Research Centre, Department of Electronic and Computer Engineering, University of Limerick, Limerick, Ireland h Department of Mechanical Engineering, Faculty of Engineering, Bayero University, Kano, 3011 Kano, Nigeria b

H I GH L IG H T S

up to date examination of the research status on dissimilar welding of titanium. • An combinations with steel, aluminium, magnesium, nickel, etc. are discussed. • Possible modification techniques to improve joint strength are highlighted. • Process mechanical properties and fracture characteristics are reviewed. • Microstructure, • Provide current progress on the issues and challenges of dissimilar welding.

A R T I C LE I N FO

A B S T R A C T

Keywords: Laser welding Titanium Dissimilar welds Mechanical properties Defects Mechanisms

Since its inception, laser beam welding as a high-quality fusion joining process has ascertained itself as an established and state of art technology exhibiting tremendous growth in a broad range of industries. This article provides a current state of understanding and detailed review of laser welding of titanium (Ti) alloys with corresponding dissimilar counterparts including steel, aluminium, magnesium, nickel, niobium, copper, etc. Particular emphasis is placed on the influence of critical processing parameters on the metallurgical features, tensile strength, hardness variation, percentage elongation and residual stress. Process modifications to improve dissimilar laser weldability by virtue of techniques such as laser offsetting, split beam, welding-brazing, hybrid welding and materials modifications by means of the introduction of single or multiple interlayers, fillers and pre-cut grooves are exploited. Detailed and comprehensive investigations on the phenomena governing the formation and distribution of the intermetallic phase, material flow mechanisms, their relations with laser parameters and their corresponding impact on the microstructural, geometrical and mechanical aspects of the welds are thoroughly examined. The critical issues related to the evolution of defects and the corresponding remedial measures applied are explored and the characteristics of fracture features reported in the literature are summarised in thematic tables. The purpose of this review is tantamount to emphasise the benefits and the growing trend of laser welding of Ti alloys in the academic sector to better exploit the process in the industry so that the applications are explored to a greater extent.

1. Introduction Titanium (Ti: atomic number 22) is a lustrous transition element



and is the 9th most abundant element on earth’s crust having a specific gravity of 5.54 g/cm3 and a density of 4.506 g/cm3 [1]. Ti is the 4th most widely available structural material after aluminium (Al), iron

Corresponding author. E-mail address: [email protected] (M.M. Quazi).

https://doi.org/10.1016/j.optlastec.2020.106090 Received 19 August 2019; Received in revised form 1 January 2020; Accepted 20 January 2020 0030-3992/ © 2020 Elsevier Ltd. All rights reserved.

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Fig. 1. Number of published documents rising throughout recent years.

plasma welding, etc. These welding techniques produce larger distortions, wider heat affected zones (HAZ) and brittle microstructures that coexist with significant residual stresses [9,10]. These implications limit the use of such conventional welding methods. Alternatively, a more desirable, well-proven and highly efficient welding technique that promotes the production of high-quality welds with sublime service properties emanates from photon-based manufacturing technology that is known as laser welding. Laser welding offers potential advantages including higher processing speeds with rapid start and stop capability, high energy density, room temperature and atmospheric pressure weldability, ease of workpiece handability, greater accuracy, minimised contamination, energy efficiency, process flexibility and narrow heat affected zones (HAZ) with subsequently lower distortion [11]. These features have made laser welding by far the most studied and presently applied processing technique. Recently, the number of publications associated with laser welding and joining of Ti-based alloys have rapidly intensified. A keyword search of “titanium” or “Ti” and “laser welding” on the science citation index expanded (SCIE) database reveals 494 documents comprising of 400 journal articles, 114 proceedings papers and 2 book chapter that have clearly exhibited a rising trend. Amongst the 494 documents available, 455 documents have been published in the recent years (1999–2019) as shown in Fig. 1. The scope of laser welding is not limited to the field of materials science only, hence, the scope of Tibased laser welds can be classified according to categories defined in the web of science database. Fig. 2 identifies a treemap for the top 15 principal categories demonstrating that apart from materials science and engineering, the Ti-based laser welds are also well-established in the field of biomedical engineering and dentistry, signifying their widespread applicability. When it comes to earlier available literature reviews, there are a few review articles that have discussed laser weld monitoring [12], laser hybrid welding [13] or laser welding under vacuum [14] while some have focussed only on the prospects for automotive applications [15]. Other work has been dedicated to specific materials of choice such as Nitinol [16,17] and bulk metallic glass [18], copper [19] and light metallic alloys of Mg [20], Al [21] and Al-Li [22]. To the extent of the authors’ knowledge, specific reviews on laser welding of titanium and its alloys, which has significant commercial importance, has not been covered yet. With the growing trend of the journal articles’ publication, it becomes mandatory to carry out a detailed literature review to investigate the role of laser in welding of Ti alloys with dissimilar engineering alloys in conjunction with the corresponding features such as changes in the microstructure, phase evolution and mechanical

(Fe) and magnesium (Mg) and is relatively light when compared with conventionally employed engineering alloys such as stainless steel (SS), nickel (Ni), cobalt (Co), etc. Moreover, Ti is as strong as commonly employed steels but is much less dense and its high melting point (1670 °C) renders it usable in high-temperature applications without creeping up to 550 °C [2]. For instance, the specific strength of β-Ti alloys is 260 kN m/kg that is almost 4 times higher than stainless steel 304, 2.2 times greater than AA7075-T6 and 1.72 times higher than Inconel X-750. Apart from strength, Ti alloys, in particular, are the only alloys that have attested excellence in all the essential mechanical properties including stiffness, fatigue life, strength, impact resistance, biocompatibility and corrosion. However, their usage has always been restricted due to the higher cost of the Ti-metal itself that is particularly associated with the difficult extraction process. With the advancement in extraction metallurgy [3,4], the cost of Ti in world markets has decreased from 21 $/kg in 2005 to that of 4.5 $/kg in 2017 and reached its lowest point in 2009 (2.5 $/kg) [5]. Given the availability of Ti base material at a considerable price, it becomes quite imperative that the potential of this metallic alloy is further explored. The microstructure of pure Ti is α phase with HCP (hexagonally close-packed) and once heated above 882 °C, it undergoes allotropic phase transformation into the ductile BCC (body centred cubic) β phase comprising of a greater number of slip systems [6]. This gives Ti a major advantage because properties of Ti are highly dependent on the heat treatment and subsequently is affected by the heating and cooling cycles imparted during laser welding. The alloying of Ti can enhance the secondary properties, thermo-mechanical processing, strengthening via heat treatment and microstructural changes, etc [7]. Ti can be alloyed based on the ability of the alloying element to stabilise the primary α or β phase depending on the atomic radius. For instance, solute elements of Al with an atomic radius in the range of 0.85–1.15 Å diffuse substitutionally by replacing one of the bulk atoms in the lattice. Whereas, atoms with radii lower than 0.59 occupy interstitial sites between the larger solvent atoms. Stabilising elements are added to prevent subsequent phase changes, increase or decrease the β-transus temperature range and enhance mechanical properties through heat treatment. The α phase is stabilised by various elements such as Al, O, N, Ga, etc., while the β phase can be stabilised by Mo, V, W and Ta. Accordingly, the Ti-alloys are classified as α or near α, α + β,β or near β alloys. Production of Ti structures requires employing processes that involve welding and joining methods that are frequently accomplished by fusion mechanism based on welding techniques [8]. Commonly employed welding techniques are tungsten inert gas, metal inert gas, 2

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Fig. 2. Treemap of publications based on the web of science categories.

evaluated by scanning electron microscopy attached to an energy dispersive X-ray spectroscopy to give some primitive results which provide analysis of the chemical, elemental and phase composition. However, to view the microstructure in detail including all features and to specify the phase and investigate the crystallographic orientation, high resolution transmission electron microscopy (HRTEM) is preferred. HRTEM likewise requires the samples to be cut in very thin sections and afterward electro-polished. X-ray diffraction (XRD) has been employed specifically to detect the primary and IMC phases present in the fusion zone (FZ) and interfaces for both welded and fractured surfaces. Furthermore, XRD is used to acquire the residual stress distribution while in-situ high energy XRD investigations can also show the lattice parameters of welded joints and depending on thermal cycles it can show stacked diffraction patterns [23]. The diffusion and accurate elemental distribution during melt pool convection and the inclusion of impurities such as C, O and N are more recognisable by a relatively reliable and rarely employed electron probe X-ray microanalyser (EPMA) method. Electron backscattering diffraction (EBSD) which is a relatively new applied technique shows the grain size, high angles grain boundaries, misorientation angles and the direction of the grain growth relative to the direction of the heat flow by means of inverse pole maps, grain boundary maps and unique grain colour maps. However, sample preparation for EBSD requires stringent measures for polishing [42]. For the sake of retrieving chemical bond formation data and understanding the bond mechanism of Ti polymer-based joints, X-ray photoelectron spectroscopy (XPS) has been harnessed extensively in the recent years. Micro and Nano level surface topographic measurements through atomic force microscopy have not been carried out for the welded seams or fractured surfaces. In-situ temperature measurements during the welding process have been taken using infrared thermography and thermocouples, as this information is crucial for evaluating the peak temperature and cooling rates over a course of time. Measurements of surface roughness as well as profiles of the seams and surface have been carried out by a roughness tester. The physical characteristics and weld integrity including cracks and defects are gauged by dye penetrant inspection and leak testing. The mechanical characteristics of welded joints are extremely important and require careful examination of properties such as hardness, modulus of elasticity, fatigue, tensile, flexural, torsional and fracture strength as well as residual stresses. To obtain information pertaining to grain size, image processing programs such as ImageJ have been utilised to estimate the mean grain size distribution through the linear intercept method [43].

performance. The review includes the associated problems and difficulties, the remedial measures taken and the gaps which are required to be foreseen for future scientists to carry on the research. The paper also intends to explore and highlight the current and potential applications with scientific evidence in favour of employing laser welding for broader industrial applications. 2. Review of laser welding techniques for Ti-alloys Laser welding of Ti and its alloys has been carried out in similar and in dissimilar material combinations. However, to reduce the scope of this review, limited topics for similar welding of Ti alloys have been considered. In general, efforts have been undertaken by researchers to investigate and optimise the effect of processing parameters on different types of Ti alloys. Fundamental studies that are essential to examine the keyhole formation mechanism, porosity development, heat flow and geometrical features of the weld have been accompanied by both experimental and simulated formulations. Similarly, dissimilar welding studies have been dissected based on different techniques employed to enhance the joint strength and suppress the intermetallic compounds (IMC) formation. Hence, researchers have focussed on an alternative modification of laser processing methods like laser offsetting, welding brazing and hybrid welding techniques and have made changes in the materials system by either employing filler materials or by adding single or multiple interlayers. Some studies have also concentrated on investigating the effect of joint type, groove shape and groove angle on the joint efficiency. Different characterisation techniques and standards used to evaluate the mechanical performance along with the general applications of Ti-based laser welds are summarised systematically in the following sections. This is accompanied by a detailed review of dissimilar welding of Ti with its dissimilar counterparts. 2.1. Characterisation techniques Researchers have employed different techniques to characterise the material and mechanical aspects of Ti-based welded joints. A summary of the techniques and methods to evaluate different aspects of the welded joints are categorised and outlined in Table 1. After performing the experimental work, the samples are prepared for either microstructural examination or for further mechanical testing. Preliminary examinations to reveal the microstructure are carried out by optical light microscopy, and special tint filters are used to obtain colour-based grain maps. The weld geometrical and microstructural features are 3

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Table 1 Characterisation techniques, mechanical testing and standards employed. Techniques Materials Characterization Optical Microscopy, Polarized and sensitive tint filter [24] Field emission (FE) scanning electron microscopy (SEM) [25] Electron probe X-ray micro analyser (EPMA) [26] X-Ray photoelectron spectroscopy (XPS) [27] Differential scanning calorimetry (DSC) [28] Transmission electron microscopy (TEM) [29] X-Ray diffraction (XRD) [30] Energy dispersive X-ray spectroscopy (EDS) [31] Focused Ion beam milling (FIB) [32] Electron backscattering diffraction (EBSD) [26] Mechanical Testing & Standards V-notched Charpy Impact Testing [33] Bending test [33] Hole Drill strain measurement method [24] Ultimate Tensile Strength Optical profilo-metre (3D), White light interferometer Vickers & Knoops and Nano-indentation [34] Other properties measured Corrosion Electrochemical impedance spectroscopy (EIS), potentiodynamic polarization Ion release test & inductively coupled plasma-optical emission spectrometer Non-destructive Testing Infra-Red Thermography, Thermocouple [35] X-Ray radiography [36] Pin photodiode sensors High-speed Imaging (HIS) [38] Computer Tomography [39,40] Photo-elastic method, Dye penetrant [41]

Usage

Analysis

Extensive Extensive Recent Recent Rare Medium Extensive Rare Rare Recent

Penetration depth [25], weld geometry, colour segregated macrostructure Microstructure, Defects, fractography, grain size, seam Elemental composition and distribution Identify chemical species present on the dissimilar metals Phase transformation order, Glass transition and melt temperature Microstructure, grain size, dislocation, orientation, Diffraction Pattern Phase information, residual strain, crystallographic information Elemental distribution TEM samples preparation and microstructure by sectioning the thickness Grain size and misorientation angles

Rare Rare Rare Extensive Rare Extensive

Impact Energy (J), Fracture characteristics Flexural strength Residual Stresses Load displacement relationship, Modulus of elasticity, fracture load Surface profile, Roughness, humping pattern, seam (peaks and valleys) Hardness distribution for WZs, stress intensity factor Residual stress, torsional strength, dynamic fatigue testing, flexural & fatigue strength, wettability

Rare

Measure Ecorr as corrosion potential, total impedance and the polarization charge-transfer resistance (RCT) Metallic ions released and their concentration

Rare

Medium Medium Rare Medium Rare Rare

Predictor of final weld quality, noncontact high-resolution surface temperature monitoring, Tmax, Tavg Measure percentage porosity, Internal defects [37] In-process monitoring signals Observe weld seam, wetting, Crack initiation and propagation behaviour Weld defects and cracks Stress Distribution, Surface Irregularities

welding, but they are devised to cover a broad range of fusion welding techniques. For instance, the American Welding Society (AWS) standard, AWS D17.1:2001 [64] has been employed to the laser beam welding of Ti–6Al–4V although the standard was meant for general purpose fusion welding. Similarly, specific standards for laser beam welding of metallic materials for aircraft applications contained in the European Standard BS EN 4678:2011 complies with the welding criteria of Ti-based welded joints. That research work has been specified for aerospace applications that require extremely stringent criteria for weld inspection and quality standards [65]. There is a need for authors to refer to guidelines that are specifically practiced for Ti-alloys while explaining the use of materials, fillers, workshop practices, equipment, process, repair of defects, etc., that have been described in details in AWS G2.4/G2.4 M:2007 [66]. Similarly, AWS also provides structural welding codes (D1.9/D1.9 M:2015 and A5.16/A5.16 M:2013) for Ti that can be availed for following research practices [67,68]. Grinding and polishing the welded samples and the subsequent image analysis details and specifications are available in this Ref. [69]. In regards to assessing the fatigue test growth rates for welded specimens, authors have followed ASTM E647 standards [70,71]. Hardness evaluation by employing Rockwell hardness tester requires ASTM-E18 guidelines for a special purpose and non-ferrous alloy materials provided the hardness values of the metal exceed 20HRC [72,73]. According to the ASTM-E8M04 and ASTM E3-13a specifications, tensile samples are prepared from the welded section and are tested at room temperature [74]. Some of the commonly employed standards are ASTM e3-11 (metallographic specimen preparation) [75], ASTM E38411e1 [76] (micro-hardness measurements), ASTM G1 (corrosion rates) [77], ISO 10271 (electrochemical test) [78], ISO 17642-3:2005 (E) (destructive weld tests) [79], HRN EN ISO 14 175: 2008 (shielding gas), B4.0 (weld mechanical testing) [80], EN 10002-1 (tensile testing) [81], ASTM E23 (V-notch Charpy impact test) [82] and ASTM E647 (fatigue crack growth rates) [70].

2.2. Surface treatments Since Ti is sensitive and reactive at elevated temperatures, great care is required before commencing of the welding process. The preparation of the surface at the onset of the welding process can be a deciding factor for the corresponding weld bead geometry, inclusions and absorption of the laser beam to generate a keyhole. Generally, treatments like sandblasting, chemical cleaning and grinding are employed, whereas black paint and graphite coatings have been ascertained to have deteriorating effects on strength [44]. To remove the burrs, after post casting or cold working, austenitic stainless-steel wire brush can be used [45]. Laser cleaning before starting the welding operation can subsequently improve the surface characteristics and enhance the weldability [46]. Chemical cleaning is able to remove contaminants and oxides, and eventually enhancing the laser beam absorption. Some authors have reported pickling in a solution of (HF:HNO3:H2O = 1:4:5) [47], cleaning with 37% phosphoric acid [48] and acetone rinsing [49] followed by ultrasonic cleaning to remove the surface oxides before welding. Drying in an oven to remove any moisture may also affect the process. For revealing the microstructure, metallographic treatments depending on the kind of material investigated have been used. Weight loss and potentiodynamic polarisation examinations can help obtaining useful knowledge on the corrosion resistance of the welds in various corrosive environments. Some of the commonly used chemical etching formulas employed for similar and dissimilar metallic combinations and corrosion mediums are tabulated in Table 2. 2.3. Standards and specifications Welding and mechanical testing that are carried out should conform to numerous well established and recognised industry standards. However, welding standards are not specifically developed for laser 4

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Table 2 Chemical solutions used for evaluating the microstructure, corrosion characteristics and cleaning of the sample surface. Techniques

Analysis

Etching CP-Ti NiTi [51] Ti–15V–3Cr–3Al–3Sn/Ti–6Al–4V [25] TA0 [52] Ti-6Al-7Nb alloy and CP Ti [53] Ti-6Al-4V, CP-Ti [36,54] AISI SS304 [55] Ti-6Al-4V:SS316L Ti-6Al-4V:AZ31B::SS304L [56] Ti6Al-4V:AZ31B 5A06Al:Ti6Al4V [57] CP-Ti/Q235B

Revealing the microstructure Weck’s reagent (5 g NH4F HF, 0.5 mL HCl conc. in 100 mL), aqueous solution) and Kroll’s reagent (1–3 mL HF conc., 2–6 mL HNO3 conc. in 100 mL aqueous solution), Kroll solution [50] Acidic mixture (17 mL of HF + 33 mL of HNO3 + 50 mL of H2O) 10 mL of lactic acid, 5 mL of hydrofluoric acid, and 30 mL of nitric acid Acidic solution (HF + HNO3 + H2O) for 2 s 32 mL H2O, 15 mL 60% HNO3 and 3 mL 46% HF Kroll reagent 3% HF, 6% HNO3, 91% distilled water for 30 s Reagent HCl: HNO3 (3:1) for 25–30 s (5 g) FeCl3, (5 mL) HCl, and (100 mL) ethanol 2 g picric acid, 10 mL acetic acid, 10 mL water and 100 mL ethanol 4.2 g picric acid, 10 mL acetic acid, 10 mL water and 100 mL ethanol for 20 s 1.5% HCl, 2.5% HNO3, and 95% H2O 10 mL of HF, 5 mL of HNO3, and 85 mL of H2O (CP-Ti side)/4 mL of HNO3 and 96 mL of alcohol (Q235B side)

Cleaning Solution Acidic Solution [58] Cleaning Solution Chloride Solution

Removing the oxide layers 4 mL hydrofluoric acid, 10 mL nitric acid, and 86 mL distilled water Acetone and ethanol 0.9% NaCl solution [59]

Evaluating Corrosion CP-Ti & Ti6Al-4V [60] NiTi [61] CP-Ti [62] CP-Ti [63]

Mediums Air, synthetic saliva and fluoride synthetic saliva Hanks solution 3.5% sodium chloride (NaCl) and 0.1% phosphoric acid (H3PO4) Ringer's solution

Laser beam dissimilar welding of Ti alloys is primarily important in structural applications involving aircraft engine (gas turbine blades, nacelle centre-beam frames, and large bulkheads). Pneumatic system of aircraft uses welded Ti seams for bleeding hot air from the engines [96]. In airplane fuselage fabrication, the wings that are made of Ti alloy could be subsequently joined with Al fuselage thus signifying the applications of dissimilar welding [97]. Welded tubes of Ti-3Al-2.5V are extensively employed in fuel transmission and hydraulic lines due to their excellent cold formability [46]. The growing trend of lasers in manufacturing and biomaterials processing can be realised in the essence of patents that have been filed or accepted for the fabrication of various devices and materials such as micro-electromechanical components [98], stents [99], prosthetic joints [100,101], resin film precision biomedical devices [102], implants [103,104], biosensors [105], batteries [106,107], and other medical instruments [108]. Thin foils of Al are required to be welded to Ti for batteries and heart pacemakers [109]. Sealing of implantable devices such as ferrules can be carried out with lasers without damaging the electronic packaging constituents [110]. Similarly, Ti-based orthodontic articles that comprise of a bracket body and a pad are typically welded [111]. Hip implants comprising of steps require porous mesh surface to be attached by means of welds [112]. It is quite imperative that with the advancement of photon-based materials’ processing techniques in the recent past, likewise its future has been envisioned for numerous biomedical applications [113]. For instance, laser welding can be employed in joining free ends of distal and proximal struts to the collar of a NiTi endoprosthetic device known as stent [114]. Welding arrangement system to weld transparent silica-containing materials (Corning GlassTm) with Ti-based materials was reported much earlier in 1982 [115]. Sporting equipment such as golf clubs that require metal pieces to be on the outside surface of the crown member are also fabricated [116,117].

2.4. Typical applications of Ti-based laser welds Prominent applications that are associated with laser welding of Ti alloys (Fig. 3) have embodied the manufacturing of biomedical devices, orthopaedic and prosthetic implants, sports equipment, aerospace structural and engine modules in addition to critical petrochemical structures, etc. Fig. 3(a) shows nomenclature of typical sections of a dental framework and Fig. 3(b) presents the repair of framework carried out by laser welding [83]. In another example (Fig. 3(c)) laser welding of Ti bars to Ti-implants is carried out as an ex vivo split mouth in pig jaws and thermal camera image was captured to determine the temperature distribution [84]. Similarly, laser welding technique is employed to repair and fabricate an orthodontic mini-implant. The head part was milled to correct the angle of C-implant (Fig. 3(d)) and the lingual part of the customised abutment was fabricated in lost-wax technique (Fig. 3(e)). Thereafter, these ready-made head parts of the Cimplant were subsequently laser welded with customised lingual part of the abutment as presented in Fig. 3(f) [85]. Likewise, laser welding of a clasp on a removable partial denture [86] is shown in Fig. 3(g) and laser welded dental framework is presented in Fig. 3(h) [87]. Laser welding is also required in the fabrication of eyelet-to-ferrule of braze case assembly for Functional Electrical Battery Powered Micro-stimulator (FEBPM) as exhibited in Fig. 3(i) [88]. Apart from biomedical applications, laser welding of seamless tubular products (Fig. 3(j)) made of Ti and Ti-alloys is gaining attention in the field of aerospace, marine, chemical industry, energy and transportation technology [89]. Fig. 3(k) shows the passenger seat track as a laser beam welded dissimilar joint conceptual design (AIRBUS) [90]. These applications have been perceived because of the higher specific strength, fatigue and creep resistance, low elastic modulus followed by good biocompatibility and excellent corrosion resistance [2]. Laser welds are involved in the fabrication of thick-walled tubes made from Ti alloys containing platinium for enhanced corrosion resistance for applications involving heat exchanger tubing [92]. Inventions are related to the use of filler materials to obtain a predetermined composition in the weld pool resulting in the enhancement of the weldability of dissimilar Ti-Fe metal, as presented by Peter et al. [93]. Similarly, to retain the weld strength of Ti-Ti alloys Kakimi et al. [94] proposed a Ti (0–3 wt% Al) filler material. Likewise, metal-ceramic based composite joints were developed by William et al. [95].

3. Laser welding of Ti and its alloys 3.1. Process overview Laser welding permits the use of a wide range of parameters that allow for precise control of thermal input which was not previously achievable [118]. For instance, laser beam welding allows higher 5

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Fig. 3. Applications of laser welding in (a, b)) dental framework [91], (c) Ti-implants [84], (d, e and f) orthodontic mini-implant [85], (g) removable partial denture [86], (h) dental framework [87], (i) micro-stimulator [88], (j) seamless tubular products [89] (k) the passenger seat track [90].

materials. Some of the important parameters that are required to be considered can be classified accordingly;

welding speeds when compared to other conventional fusing welding techniques. The rapidness of laser welding along with exceptional productivity and repeatability of the process makes it suitable for industries that require automation. The formation of the weld seam and the quality are strongly influenced by the processing parameters and more so in the case of pulsed laser parameters. For instance, in dissimilar welding of materials, the melt pool size in FZ is unsymmetrical due to the difference in thermal properties. The melt pool tends to be larger for a material that has relatively lesser thermal conductivity than its counterpart. In addition to the thermos-physical properties, the absorption characteristics of the material at the incident laser beam wavelength can affect the process efficiency, keyhole stability, penetration depth and may cause defects formation. Fig. 4 presents a comparison of the energy absorptivity of various metallic materials at room temperature, depicting that Ti exhibits reasonably good absorption characteristics. However, the absorptivity of some metals increases when melting temperatures are reached and their reference values are stated in the earlier work of Xie et al. [119]. Hence the welding characteristics become dependent on the thermos-physical characteristics of the materials. Some of these basic characteristics of Ti and its alloys are listed in Table 3 and compared to conventionally employed structural

1- Laser related: active medium, pulsed or continuous beam, pulsed duration, pulse frequency, scanning speed, power density, peak power, incident angle, power average, spot size, position of the beam, etc. 2- Material related: thermo-physical properties, alloying composition, heat treatment, geometrical dimensions, microstructure and joint configuration. 3- Shielding gas: gas flow-rate, composition, configuration and the angle of incidence. 4- Filler material: composition, geometrical dimensions and thermophysical properties. The difference in conductivity of the dissimilar materials directly influences the symmetry of the weld, composition and asymmetric heat transport. A previous study has scanned the laser beam normal to the rolling direction of the as-received Ti-plates [25]. The gap between the materials to be joined and the joint configuration along with the end shape are the parameters that can derive the weld pool geometry. The 6

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Fig. 4. Absorptivity of various metals at room temperature when subjected to Nd: YAG laser (1064 nm) interaction [119,121] (Absorption data for Mg and Co available for a wavelength of 1000 nm taken from Ref. [122]).

addition to the scanning speed and focal distance. With control of a greater number of parameters, PW laser welding exhibits the advantages of low heat input, shorter welding cycle, higher accuracy of energy input position and is able to joint small components [123]. The melt pool characteristics, resulting phase constituents, mechanical properties of the weld and the failure modes are strongly affected by the pulse profiles that are employed in the process. The pulsed input has been able to effectively reduce the IMC formation and enhance the homogeneity of the stainless steel (SS)/Ti-based weld [124]. In comparison to the rectangular pulse profile, the ramp down profile transferred lesser energy which reduced the convective fluid flow, thus producing a smaller weld size. The lesser Marangoni flow results in a lesser degree of intermixing of the two phases that reduce the likelihood

laser welding operation is entirely a complex cloud of parameters that consequently derive the shape and properties of the molten pool that evolves the welded region. 3.2. Continuous laser vs pulsed laser Laser welding can be performed either in a continuous wave (CW) mode or pulsed wave (PW) mode. PW mode provides better control, smoother seam characteristics while producing deeper penetration. The reason is that in CW mode a smaller number of parameters can be controlled such as scanning speed, laser power and stand-off distance. While, in the case of PW mode, more parameters can be controlled, such as pulse power, pulse duration, pulse shape, pulse repetition rate in

Table 3 Thermo-physical and tensile strength of different metals and alloys that are commonly welded with Ti [120] Material

Melting point (°C)

Boiling point (°C)

Density (g/ cm3)

Specific heat (kJ/kg K)

Thermal conductivity (W/m K)

Coefficient of linear thermal expansion [m/(m K)e-6]

Latent heat of fusion (kJ/kg)

Tensile strength (MPa)

Titanium Ti-6Al-4V Aluminium Magnesium Iron Niobium Inconel 600 Vanadium Tantalum (Cold Worked)

1668 1655 660 650 1536 2468 1300 1910 2850

3287 3315 2470 1091 2862 4900 2917 3407 6000

4.50 4.43 2.385 1.59 7.015 8.55 8.91 6.11 16.65

0.520 0.61 1.08 1.36 0.795 0.27 0.435 0.489 0.14

7 6.7 94.03 78 38 53.7 11.4 31 54.4

8.6 8.6 22 26 12 7.1 13 8.33 6.5

419 290 398 368 272 285 272 410 172

960 1100 310 280 540 275 272 800 900

7

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of brittle IMC formation that can finally affect the fracture mode. The penetration depth and the weld pool width are also a function of the pulse width and peak power as exhibited in welds of Ti-6Al-4V [125]. The stress concentration factor was found to be higher for CW joints of Ti-2Al-1.5Mn at the edges of the concavity of the weld seams when compared with PW joints [126]. Lower pulse energy produces a shallow and narrow melt pool in Ti-6Al-4V that is relatively smoother than high pulse energy seams [127]. Pulse overlapping is also a factor that contributes to the set of pulses that hit the area of a spot. A combination of process parameters such as pulse energy, pulse duration, pulse repetition rate and travel speed determine the overlapping factor values. By increasing the overlapping, the re-melt and re-solidification quantities increase, and the pre-heat region for the next pulses expands [128]. 3.3. Effect of heat input The heat input represents the energy supplied to the workpiece. The heat input is one of the most influential and controllable parameters that can alter the chemical composition of the melt pool, the geometrical constraints, the degree of dilution and the formation of defects. For instance, increasing the amount of heat input causes changes in the seam shape [129]. The changes in the weld pool geometry are dependent on the cooling rate which is inversely proportional to the square of the weld pool length [130]. By restricting the size and extent of the melt pool the IMC phases can be restricted as the peak temperature and solidification time are altered. The Maragoni fluid flow that causes flow reversal is also observed when temperature gradients are instantaneous causing a narrower FZ and encouraging the formation of IMC. Rapid cooling can be the basis for thermal strains and crack initiations but metallic alloys that are not prone to cracking can reap the advantage of high cooling rates in the form of microstructural refinements leading to a rise in the hardness of FZ region. However, some microstructures are prone to segregation in the weld pool due to non-equilibrium solidification in the weld pool of dissimilar materials. The laser power can also affect the convective flow also termed as Marangoni effect wherein the flow of the fluid is governed by the spatial gradient of the surface tension at the melt pool surface owing to the variation of its local temperature and density [131].

Fig. 5. Effect of changing the welding speed on the grain size. The inserts show the change in the EBSD (inverse pole figure + grain boundary) images of the fusion zone [133].

the most common shielding gases that have been employed [134]. During the shielded welding process, the plasma above the melt pool comprises noble unreactive shielding gas and Ti ions which produce a relatively stable flux [96]. Helium is preferred as it has a lesser specific weight than air and Ar [43] that is beneficial to shield the bottom part of the weld. It is quite important to use effective shielding at both the face and root of the weld to prevent the porosity from entering from entry points at the top and bottom regions [56]. A wider gap between the joints can produce undercut defects while a narrower joint can result in underfilling at the root section. These defects are also arising due to plasma instability owing to the entrainment of the contamination from the air. The inclusion of air makes a variation in the intensity of plasma and blocks the incident laser radiation, thus causing porosities, cracks and concavity [96]. The flow rate of the shielding gas needs careful selection as excessively higher flowrates tend to decrease the UTS of the Ti-joints [135] while ineffective nozzle angle and stand-off distance also matter [136]. Laser welding of Ti under vacuum has rarely been carried out but a preliminary investigation has shown that residual oxygen is not present and penetration depth as high as 40 mm is obtainable, so fulfilling the DIN EN ISO 13919 requirements [137]. The solubility of hydrogen is very high in the β phase because the tetrahedral holes are larger in BCC than HCP but it can be removed through annealing in a vacuum [6].

3.4. Effect of scanning speed Reducing the scanning speed leads to higher heat input that provides coarser grain size and lesser grain boundaries in the weld. Besides, higher scanning speed generally leads to higher cooling rates that can alter the keyhole geometry, melt pool stability and beam absorption. The combination of welding speed and heat input is also able to suppress the formation of IMCs in certain cases [132]. A rise in the welding speed can alter the shape of the weld pool. Parallel shaped weld bead is formed at a lower speed that transforms into wedge shape at a higher scanning rate due to change in the surface tension [133]. An increase in the scanning speed for grade 2 Ti caused a decrease in grain size (Fig. 5) due to higher cooling rates that resulted in higher weld strength. The microstructure evolution based on the misorientation angle in the FZ was characterised by a greater number of low angled grain boundaries (formed as α substructure) and smaller fraction of higher angled grain boundaries (formed as α originating prior β-grains) with dislocation tangles. At higher power, the effect of scanning speed on grain size and hardness becomes more noticeable [26]. It should be convenient here to remind that both power and welding rate (scanning speed) contribute to the net power density applied to the welding zone (energy proportional to power/welding rate).

3.6. Keyhole vs conduction modes The formation of the solid/liquid and liquid/gas interface is of utmost importance to academia as it defines whether melting occurs as a surface or keyhole. In general, the energy absorbed through the laser beam takes place through either keyhole or conduction mode via interaction with materials. The power density for the conduction mode is lower than 103 W/cm2 but it is considerably higher for the keyhole mode, being in the range of 105–107 W/cm2 [138]. The keyhole mode of welding process allows the laser beam to produce welds that are deep and narrow. These modes mainly govern the amount of laser beam energy absorbed by the target either through a wider melt pool surfaceconduction mode (20%) or by a deeper and narrowly filled cavity described as keyhole (70–90%). These modes can significantly alter the weld characteristics. Different materials exhibit variance in the transition from conduction to the keyhole and are dependent on materials’ thermal properties (thermal conductivity, specific heat capacity, melting point and vaporisation temperature). It is known that energy absorption from laser pulse is active during heating in both the solid and liquid phases of the metal until reaching the vaporisation temperature, wherein keyhole starts to form. Keyhole creation results in a

3.5. Shielding gas Shielding gas is employed to protect the weld from various inclusions, porosities and other defects [38]. Argon (Ar) and helium (He) are 8

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Fig. 6. Comparison of liquid/solid interface during the evolution of keyhole formation between simulated model and high-speed camera imaging using 100 mm/s at 3000 W. Note that t0 refers to the initial time at which the laser contacts the weld plate [37].

For instance, the pulsed TIG (Tungsten inert gas) registered a greater spread of tensile residual stresses until about 18 mm from the weld centreline. However, laser welded samples exhibited compressive residual stresses at less than 3 mm from the centreline [144]. Furthermore, Gao et al. [10] made a comparative evaluation of angular distortion for laser welded and TIG welded Ti6Al4V. They reported a maximum angular distortion of 2.3° for laser weld and 5.5° for TIG welded specimens.

significant increase in the energy absorption through the inverse Bremsstrahlung mechanism and Fresnel reflections [139]. By employing a high-speed camera in conjunction with computational fluid dynamics, Panwisawas et al. [37] simulated and verified the keyhole formation and penetration depth at successive time intervals for Ti-6Al4V (Fig. 6). The different nature of the solid/liquid interface allows the phenomena to be captured and becomes interesting as the prediction of the fusion zone (FZ) boundary is of great interest to the industry and academia.

4. Dissimilar welding of Ti and its alloys 3.7. Residual stress

Heterogeneously assembled multi-materials are beneficial in obtaining a multi-functional advanced materials’ component that is able to perform and sustain in heterogeneous conditions, it can also reduce the specific weight and prevent waste of materials [152]. In the coming sections, mechanical and materials’ aspects of such dissimilar welding are explored.

The information pertaining to the residual stresses are of considerable importance when the structural integrity of the system is of the primary concern. Table 4 provides a summary of different techniques employed and the corresponding residual stresses measured for different Ti alloys. During laser welding, regardless of the workpiece fixture, the instantaneous melting of the material by laser and the subsequent cooling process behind the beam by shielding gas instigates thermal cycles [24]. The concurrent heating and cooling gradients initiate differential plastic flow along with thermal strains and stresses due to the phase changes that are restrained by the clamped structure and thus induces collective residual stresses in the system that could be detrimental for the buckling strength, fracture strength and fatigue strength. An unclamped system generates lesser residual stresses because a part of the strain is recovered as the plate deforms, shrinks or bends during cooling [140]. Massab et al. [24] revealed that the longitudinal and transverse stresses generated in pulsed wave mode laser welding of unclamped Ti-5Al-2.5Sn were around 100 MPa which was far lesser than the yield strength of the parent metal (950 MPa). The location of the residual stress accumulation depends on the source. For instance, stresses caused by thermal strain are substantiated towards the centreline of the weld but they shift away as the phase change complicates the stress distribution [141]. For the case of dissimilar Al-Ti welds, owing to higher deformability of Al, greater distortion is sustained in the Al side. However, if the weldments are retained in the clamp for a day, the compressive stresses eliminate the distortion because of the prolonged duration [142]. The higher the ductility of the material is, the greater will be the plastic deformation and vertical lift [142]. Lastly, age hardening is a viable approach to completely eradicate the distortion [143]. When compared with high heat input welding processes, laser based welds register a lower distortion [144].

4.1. Titanium – Titanium based joints Dissimilar joints of Ti alloys have been shown to retain the mechanical properties of the parent alloys. The primary reason is the lesser difference in thermos-physical properties and chemical compatibility that prevents the formation of IMC. Some of the laser welding parameters and their corresponding tensile fracture features are given in Table 5. It can be inferred that the three major types of laser mainly CO2, fibre and Nd: YAG have been used with beam parameter modes in the form of either continuous or pulsed wave mode. The incident applied power in the range of 0.65–4.1 kW has produced welds with sufficient hardness (Fig. 7(a)) and UTS (Fig. 7(b)) and has been protected from porosities by employing Ar gas. The welding process is sensitive and low beam energy can signify incomplete penetration while undercuts and burnt through skins could be observed at high laser power [36]. Furthermore, the angle of incidence and the incidence offset can affect the formability of defects such as less root penetration, undercut or high skin penetration for the case of T-joints configuration [36]. To enhance the mechanical properties of the joint, post weld heat treatment (PWHT) is carried out. PWHT has shown promising results in enhancing the ductility (elongation) and the hardness of the FZ, HAZ and base metal (BM). However, PWHT contributes to a very slight improvement in the UTS [25]. Similarly, dual beam welding mode exhibited higher and more homogenous distribution of the hardness in the weld zone (WZ) while enhancing the UTS (7%) and percentage 9

120 – 500

1, HAZ – 2 2.5

Ti6Al4V 162.0 J/mm

10 2.8 1.4

Not Given

Fibre, CW

Nd: YAG, PW

CO2, CW

Nd: YAG, PW

Fibre, CW

CO2, CW

CO2, CW

Ti–22Al–25Nb: TA15 [153]

TC4:TA15 [155]

Ti-22Al-25Nb: TA15 [154]

Ti–15V–3Cr–3Al–3Sn: Ti–6Al–4V [156] BTi-6431S: TA15 [58]

Ti-6Al-4V: BTi6431S [157]

Ti–6Al–4V: Ti–6Al–6V–2Sn [25,155]

Ti–22Al–27Nb: TC4 [158] 1.2

2.8

3.2

1.3

2.5–4.1

1.3

0.65

Fibre, CW

Ti-6Al-4V: CP-Ti [36]

Power (kW)

Laser Type/ Wave mode

Materials

1000

100

1600

1200

100

1200

2500

1200

3500

Scanning Speed (mm/min)

Ar

Ar

Pulse 20 ms, Spot Size 0.4 mm, offset 0–0.2 mm Spot Size 0.3 mm,

Ar

Pulse Duration 20 ms, Spot Size 0.4 mm

Ar

Ar, 15 l/min

Ar, 3–11 l/min

Other Parameters

Butt, 92%

Butt, 90%

Double sided Butt

Butt, 100%

Butt

Butt, 91.6%

Butt, 90%

Butt

Tee Joint

Joint Efficiency & Configuration

B2 phase and martensitic a′ phase

Fine α + β microstructures (basket weave)

Pre-dominant α′′ with a few α and β phase β columnar grains consisted of martensite α′ and slender acicular α Widmannstatten, Martensite α′

Coarse columnar grains, acicular α and martensite α′ B2 phase and small amount of martensite α′



Martensite α′ due to β → a′ transformation Cellular grain, B2 and O

FZ

NG



BM

NG

FZ, TA15

TA15

FZ

Fracture

2.5



318 (Von Mises) –

120

580 – – 500

– 660

– 660 700

Total residual Stress (MPa)

Dual beam mode transformed the quasicleavage feature into dimple signifying enhancement in ductility Equiaxed dimples uniformly distributed on the whole fracture surface led to ductile failure Intergranular brittle fracture with dimple gliding fracture at room temperature and ductile fracture at 550 °C Increase in PWHT temperature led to less flat facet fracture and improved ductility Ductile failure, fracture oriented at 45° with the evidence of necking. Reducing scanning speed enhances high temperature UTS PWHT at/above 593 °C, the welds were fractured by a mixture of transgranular dimples and intergranular shear –



Fracture Characteristics

195

– −90



– – 0

– 95 80 –

40 –

65 200 44

Max transverse Stress (MPa)

– 1 1 –

0 approx., FZ –

FZ Microstructure

Table 5 A summary of welding parameters for various dissimilar Ti-Ti laser joints and their corresponding microstructural and fracture features.

Neutron diffraction [151] Thermal elastic–plastic finite element simulation [151]

X-ray diffraction [150]

720

450 440 170 450

– 1 1 1, HAZ

Ti6Al4V (Argon Cooled) Ti–2Al–1.5Mn (0 N) Ti–2Al–1.5Mn (500 N) Ti-48Al-1Cr-1.5Nb-1Mn-0.2Si0.5B (As weld) Ti-48Al-1Cr-1.5Nb-1Mn-0.2Si0.5B(PWHT) CP-Ti, Grade 2 Ti6Al4V 115.7 J/mm

Finite element simulation [149]

160 455

0 approx. FZ –

CP-Ti rolled sheets Ti6Al4V (Not cooled)

11 0 1

−100 590 740

2.5 0.5 0

Ti-5Al-2.5Sn (Unclamped) Ti6Al4V Ti-6Al-4V

Location of maximum transverse stress amplitude from the weld centreline (mm)

Max longitudinal Stress (MPa)

Hole-drill strain method [24] Numerical Simulation [145] X-ray and neutron diffraction techniques, distortion [146] measurement Neutron diffraction [147] Numerical Simulation & marking and measuring method [148]

Location of maximum longitudinal stress amplitude from the weld centreline (mm)

Material

Residual Stress Measurement Technique

Table 4 Measurement techniques employed and the corresponding maximum residual stress values recorded for various Ti-alloys.

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Fig. 7. (a) Hardness distribution and (b) tensile strength values for Ti-Ti dissimilar laser welds [25,58,153–158].

microstructure is transformed from primary α and prior β phases into martensite α’ and acicular α phases. The increase in the hardness was observed as the offset was shifted from TA15 alloy towards BTi6431alloy due to an increase in martensite α′ and higher alloying content availability in the FZ [58]. The β stabilising elements such as Nb, Mo and W provide a stronger reinforcement than that of α stabilisers, so increasing hardness in HAZ for β based alloys. Li et al. [154] discussed the microstructural evolution of Ti-22Al-25Nb (α2 + B2 + O) welded with TA15 (α + β) in details. The FZ was composed of the dendritic structure growing radially. The B2 phase nucleated and acicular martensite α′ phase (Fig. 8(a)) was formed when the β phase was rapidly cooled, thereby not allowing the β to transform into α or any other phases. The HAZ of the Ti-22Al-25Nb was analysed based on the Ti-22Al-XNb phase diagram (Fig. 8(j)). At the HAZ zone, adjacent to the FZ, the peak temperature was greater than transus B2 temperature, however, the faster cooling rate suppressed the transformation of β into O/α2 causing B2 phase to dominate in this region. When the temperature decreased, the B2 + α2 phase was retained that was narrow. As the HAZ approached the BM, the peak heating temperature was alleviated to B2 + O + α2 phases (Fig. 8(c) and (d)). The changes in the microstructure of the FZ and HAZ do not affect the joint’s UTS but may reduce the ductility causing a decrease in elongation to around 60–81% [154] in some cases because of the evolution of brittle and harder martensitic α’ phase. It needs to be iterated that research work on impact strength, fatigue strength, crack initiation site, fatigue failure, cyclic strain characteristics, cycles to failure and crack propagation characteristics for dissimilar Ti-Ti joints has not been realised yet.

elongation (37%) [153]. This was caused by slower cooling rates for the case of dual beam leading to the formation of a harder O brittle phase in the relatively softer B2. This was caused by slower cooling rates for the case of dual beam welding of Ti–22Al–25Nb/TA15 dissimilar material joint, leading to the formation of a harder and brittle orthorhombic Ti2AlNb (O) brittle phase in a relatively softer Ti2AlNb (B2) phase. The O phase further restricts the dislocation movement by hindering the slip modes, so eventually improving the mechanical strength. The welds have generally failed either in a ductile manner making a 45° angle with the tensile axis or in a mixed brittle-ductile manner due to the loss of the ductility [58,154]. Xu et al. [155] showed that when the strain rates increase for the joint between TC4:TA15, the UTS also increases but the strain hardening decreases because the plastic flow mechanism is transformed from slippage dislocation to that of twin crystals. The tensile test at higher temperatures (550 °C) has shown a 30–40% reduction and 40% improvement in elongation as compared to the room temperature strength [58,157]. This reduction might seem to be significant but the second generation Ti alloys (Nb greater than 25%) remain suitable for high-temperature application of gas turbine systems (compressor portions). These alloys offer 40% weight reduction as compared to Ni alloys and exhibit higher service temperatures against Ti-Al based alloys [154]. Nevertheless, the decrease in the bearing capacity was mainly attributed to the transformation of deformation mechanism for smaller grains from that of dislocation slip into the creep. For welds with coarser grains, the size of precipitates is normally large due to the longer time available for their growth at slow cooling. [157]. In some cases, the induced porosity is inevitable, and if it is limited within a 5% range, it does not affect the bearing capacity of the joints [58]. However, in the case of double side welding of Ti-6Al-4V and BTi6431S, more porosities are evolved in the second pass as the smaller porosities in the first pass accumulate to form a large one in the second [157]. Furthermore, it is noteworthy that pulsed laser welding signifies superior mechanical properties of the (Ti-22Al-25Nb/TA15) joint owing to its discrete heating and cooling cycles [154]. The microstructure of the FZ is refined after laser welding and is found as a mixture that is based on the intermediate phase and alloying composition. Gui et al. [157] observed hard Widmanstätten structure (400 Hv) in the FZ of the overlap zone of double side laser welding between Ti-6Al-4V and BTi6431S but there were no changes reported in the HAZ for single and double pass welds. It was iterated that the UTS values of the dissimilar weld were intermediate between BTi6431S and Ti-6Al-4V due to an increase in Mo content (β-stabiliser) from 2.5 to 3.3 wt%. As presented in Fig. 7(a), the hardness of the HAZ in some cases seems to be higher than that of FZ and BM. The primary microstructures and phases in Ti alloys follow the order of micro-hardness: martensite α’ > α phase > β phase [58]. Hence, in HAZ, the

4.2. Titanium-steel joints Titanium-stainless steel (Ti-SS) welded joints are of particular interest for petrochemical and nuclear industries wherein process vessels are made of load-carrying steel base protected against environmental degradation and corrosive aqueous attack by Ti alloys. Likewise, Ti-SS sheets have pronounced prospects in seamed tubing [89], piping and vessel components [33]. Comparably, Ti welded structures are prominent in reducing the weight of the fuselage and components used in aerospace trades [49]. Joining Ti to steel is restricted by their very limited solid solubility (0.1 at. %) at ambient conditions and mutual solid solubility in liquid conditions [49]. The variation of hardness across the weld pool for different Ti-SS joints has been presented in Fig. 9. The laser-welded Ti-SS joint undergoes spontaneous cracking under mechanical load and fails in a brittle manner because the FZ comprises of layers of FeTi (600 Hv) and Fe2Ti (1000 Hv) and other IMCs depending on the alloying composition. Since autogenous welding does not produce joints with the desired strength, various techniques 11

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Fig. 8. (a) B2 phase along with martensite α′ phase in the FZ, (e) TA15 HAZ microstructure and the corresponding magnified zones (b) recrystallization zone, (f) fine grain zone, (g) coarse grain zone, (h) martensite α′ in coarse grain zone. (i) Ti-Al Phase diagram with the sketch of the HAZ phases for TA15 alloy. The optical images of HAZ on the Ti-22Al-25 Nb side (c) overview and (d) magnified view. (i) Correlation of the TA15 Ti-Al phase diagram corresponding to the obtained microstructural features and (j) T-22 at.% Al-Nb phase diagram [154].

Fig. 9. Hardness of various zones for different types of titanium steel dissimilar joints [34,49,124,160–163]. 12

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Fig. 10. Enhancement in tensile properties due to changes in the laser welding technique for titanium steel dissimilar joints [30,33,34,49,56,124,159–165].

embrittlement irrespective of the utilised cooling rate. The solid solubility of Fe in α-Ti is extremely low (0.05–0.1%) at room temperature due to their distinct crystal lattice structures. Hence, Ti/Fe IMC was formed at the interface of the lap joint between Ti-6Al-4V and 42CrMo steel while IMC thickness increased with an augmentation of heat input, proving detrimental for the joint strength [168]. Ti readily reacts with other elements to form brittle-based IMCs except for elements like Zr, V, Nb, Ta, Mo and W [165] (Fig. 11(b)). Casavola et al. [169] numerically proved that the stress concentration in Ti butt welded joints was sensitive to the defects and seam geometry. To improve the joint strength, Chen et al. [30] used an offset of 0.6 mm towards the steel side to obtain a joint with a UTS of 150 MPa. The spontaneous failure occurred without any offset for Ti-6Al-4V welded with SS201 steel but two IMC layers of uniform thickness were formed. The tensile strength decreased as the pulse energy was reduced by varying the pulse ramp down [124].

have been developed over the past decade to increase the upper limit of the achievable strength. For instance, inserts and interlayers of compatible materials like Mg, Cu, Nb and Ta have shown sufficient enhancement in weld strength. However, due to the incompatibility of either of the BM with the insert, the weld failed in a brittle manner at the IMC interface. To further enhance the strength and prevent Ti-Fe mixing, multiple layers of Ta/V/Fe have been proven to provide the highest obtainable strength equivalent to annealed V interlayer as presented in Fig. 10. Similarly, hybrid welding technique, wherein welded multiple interlayer inserts are employed, has shown the second highest strength. A summary of the laser welding parameters and their corresponding fracture properties are summarised in Table 6. As presented in Fig. 11(a), a few phases are formed including α-Ti, β-Ti, FeTi (BCC), Fe2Ti (λ-BCC) and α-Fe in the order of increasing composition of Fe. Furthermore, eutectics of various compositions are formed near the Ti-rich side that can intensely affect the plastic properties of the joints. If the steel comprises of Ni and Cr then TiCr2, NiTi2 and Ti5Fe17Cr5 are likely to be formed. However, the low joint strength is also attributed to the thermally induced stresses developed due to stress concentration generating as a result of a variation in the phase present and thermal expansion coefficient. The eutectic reaction at equilibrium consists of alternating plates of β-Ti and TiFe in a lamellar structure. The TiFe dendrites are formed in the supersaturated β-Ti(Fe) matrix and solidify in the direction of the heat flow. For a small amount of mixing for Fe-Ti pair, the β-Ti is formed when the %Mo content exceeds the 10% equivalent threshold [159].

4.2.2. Single interlayers In order to prevent the intermixing of steel and Ti, many authors have proposed to insert interlayers that can prevent IMC formation and are chemically compatible. Cu-based inserts can prevent Ti-Fe IMC formation but at the same time, a relatively less brittle Cu-Ti IMC is also produced that could degrade the ductility of the joints. Cherepanov et al. [162] confirmed that a negative defocus of the laser on Cu insert between VT1-0Ti and 12Kh18N10T steel produced an inhomogeneous molten pool characterised by porosity, whereas cracks and joint failure was observed when the focal point was positioned directly onto the Cu insert. The WZ was composed of a supersaturated solid solution of Fe, Cr, Ni and Ti in the crystal lattice of Cu. The IMC of Ti(Fe, Cr) 2, with the size range(5–200 μm), and TiCu3 (1–10 μm), were uniformly distributed while diffusion boundaries of 0.1–0.2 mm were formed at the SS and Ti sides [31]. Zhang et al. used Cu interlayer between TC4 and SS301L to produce a joint strength of 350 MPa. The WZ comprised of Cu solid solution, Cu2Ti (acicular and granular), Cu–Ti and Cu–Ti2 [34]. The Cu2Ti was formed due to the higher amount of residual liquid copper (L → Cu–Ti), which caused a peritectic reaction (L + Cu–Ti → CuTi4). Thereafter, the peritectic reaction was as follows (L + Cu4Ti3 → Cu2Ti), forming the Cu2Ti phase [170]. The hardness of the Ti-Fe IMC was lower than Ti-Cu IMC, whereas the Cu2Ti IMC was

4.2.1. Ti-SS autogenous welding The autogenous welding of Ti-SS pair has not yielded sufficient mechanical strength (35% UTS of Ti) while the brittle failure without necking occurs in the FZ or at the Ti-SS weld interface [159]. Interdendritic cracks perpendicular to the weld direction are generated as the SS side experiences higher contraction and tensile stress while some cracks appeared in the SS-weld interface contain the λ-Laves phase. The upper weld zone (WZ) consists primarily of the dendrites of β-Ti and TiFe IMC while the lower part contains fracture resistant single phase βTi(Fe). Shanmugarajan and Padmanabham [41] reported that the autogenous welding of Ti-SS dissimilar combination caused cracking and 13

14

1, 1

2, 1, 2

1, 1, 1

0.8, 0.4, 0.8

0.8, 1, 0.8

3, 1, 3

2, 1.06, 2

Ti–6Al–4V:SS201 [30]

Single Interlayer Ti–6Al–4V::AZ31B::AISI304L [56]

TC4Ti::Nb:: SS301L [134]

TC4::Cu::SS301L [34]

TC4::Nb::SS301L [164]

VT1-0 Ti::Cu::12Kh18N10 [162]

AISI316L::V:: Ti6–Al4–6V [161]

Hybrid Welding AISI321::SS:Cu:Nb:Ta:CT1-0 [49] AISI316::Cu3Si::Ti6Al4V [163]

CP-Ti:Ti::Cu::Fe:Q235B [33]

3, 0.2, 1, 0.2, 3 2, 1, 2

0.8, 0.2, 1, 0.2, 0.8 7, 1.2, 1.2, 1.2, 1

0.25 t

Autogenous dissimilar Weld CP Ti:SS304 [124]

Multiple Interlayer TC4i:Ta::V::Fe:SS301L [165]

Thickness (mm)

Materials

CO2, Continuous Fibre, continuous, Arc Welder

Fibre, Continuous

Nd: YAG pulsed

Yb:YAG, continuous

CO2, Continuous

Pulsed

Nd: YAG, Pulsed

Nd: YAG, Pulsed

1.5–2.4 1.5

0.184, 0.127 2.5–4

3, 2.5

2.35

0.164, 1.05 6.09

63–80 A

2.5

2

CO2, Continuous

Fibre, Continuous

4.5

Power (kW)

Nd:YAG, Pulsed

Laser Type

0.9 to 1.2 m/min 1.5 m/min, Focussed at edge of bevel of SS316 Side

1.5–4 mm/min, 5-degree beam angle

250 mm/min, 8–10 ms, 6 Hz

2 m/min; Focussed at both interfaces

250 mm/min, 10 ms, Focussed at Cu 100 mm/min,15 ms, 0.2 mm offset Ti 1-m/min, Focussed at interlayer

100–250 mm/min, 8–15 ms

2 m/min

100 m/sec, 7.5 HZ, 1 ms, lap joint 2 m/min

Welding Parameters

Nb, Cu and Ta Cu3Si

Ti, Cu, Fe

Ta/V/Fe

V

Cu layer and TiN + Y2O3 + Cr nano-powder

Nb

Cu

Nb

Mg

Nill

Nill

Interlayer

Ductile fracture Brittle cleavage fracture

Brittle fracture inside the FZ

Ductile with equiaxed dimples

Brittle fracture, cleavage and quasi-cleavage facets correspond to coarse Ti(Fe, Cr)2 IMC or diffusion zones Deformed V foil with the zones of brittle fracture

Brittle fracture, secondary crack zone, river pattern on the cleavage plane Brittle joint fracture

Rough area with ductile–brittle fracture, equiaxed dimple patterns & small amount of quasi-cleavages Fracture mode of thejoint was brittle fracture

Complex fracture at SS side and brittle fracture at CP Ti side Top section smooth river pattern while rougher at the bottom with cracks throughout

Fracture Characteristics

Table 6 A summary of laser welding parameters with different techniques employed for welding titanium steel dissimilar joints and their corresponding fracture characteristics.

Cu plate CuTi2 layer at Ti6Al4V

WZ

Unmelted V

Ti/V interface

Reaction layer at Nb/ SS interface Cu–Ti2 layer adjacent to TC4 Side Reaction layer at SS side Interface with the Ti alloy

SS/Weld interface

Adjacent IMC layers

WZ

Fracture Location

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Fig. 11. (a) The binary (Fe-Ti) phase diagram showing the presence of two IMCs [166], (b) solubility properties of Ti and Fe with the other elements in the binary alloy periodic system (investigated up to 1200 °C) [167].

the SS side remained un-melted with the emanation of a reaction layer. The reaction layer was composed of Nb/Fe7Nb6/Fe2Nb + α-Fe/α-Fe/SS phase as Fe has a faster diffusion rate than Nb. When the same author employed double pass welding that required melting the Nb-SS side, γFe and Fe2Nb IMCs were formed in the WZ, which deteriorated the UTS to 170 MPa while the fracture occurred at the Nb-SS interface [134]. The subsequent failure was attributed to the hardness difference in the formed phase that contributed to crack development and residual

harder than the latter. In order to prevent the mixing of Ti and Fe, the authors proposed laser offsetting of the Nb interlayer that has a higher melting point than the parent materials [164]. Hence, the weld was characterised by fusion welding at the Ti-Nb side and inter-atomic diffusion bonding accompanied by eutectic reaction at the Nb-SS side, thereby, creating a hybrid joint having an appreciable UTS of 370 MPa. As set forth in Fig. 12(a), the Nb at the Ti side was mixed without any IMC being formed while

Fig. 12. Melt zone of single interlayers (a) Nb [164], (b) V [161] and (c) Mg (AA31B) [56]. The melt zone for multiple interlayers (d) Ti/Cu/Fe (e) cracks present at the interface, (f) mixing of Cu/Ti/Fe, (g) bending load vs deflection behaviour [33], (h) Ta/V/Fe [171] and (i) Ti/Nb/Cu/SS [49]. 15

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Fig. 13. Bead formation mechanism, (a) the distribution of melted materials and possible flows in the molten pool, (b) insufficient heat input and fast welding speed, (c) moderate heat input and welding speed, (d) sufficient heat input and slow welding speed [163].

stresses. The laser beam focussed on the AZ31B interlayer produced weld efficiency of 94%. The joining occurred due to atomic diffusion at the Ti-6Al-4V and SS304L sides. The melt zone of the joint is shown in Fig. 12(b). The Ti side also experienced reactive diffusion due to the formation of Mg17Al12 IMC because Ti retains more heat due to its lower thermal conductivity and has Al as an IMC promoter [56]. Since Mg does not dissolve or react with Fe and Ti in the liquid state, the WZ was characterised by two micrometre scaled thin diffusion interfacial layers. As the laser power was raised, the fracture shifted from the SS/ weld side to the Ti/weld side due to the augmentation in the IMC brittle phase. One pass of welding on Ti-6Al-4V/SS316L joint with a laser beam focussed on the V interlayer causes intermixing of Fe, V and Ti, so instigating ductile to brittle transition of V alloys [161]. However, double pass laser welding causes annealing and subsequently fracture at the Ti-V interface, thus producing a joint with 92% of the V interlayer. The SEM etched microstructure of Ti-6Al-4V and SS316L joint with V interlayer in a 2 pass welding mode with the magnified insets showing annealed grain size of un-melted V is in Fig. 12(c) [161]. About 89% of the initial foil thickness is retained and the annealed zone along with grain size was larger at the Ti side than the SS side due to lower thermal diffusivity of Ti-6Al-4V (2.106 m2 s−1) compared to steel (4.106 m2 s−1) and vanadium (10.106 m2 s−1) [161].

filler required and a decrease in residual stresses. Ning et al. [33] employed multi-pass laser butt welding to join explosively welded CP-Ti/ Q235 bimetallic sheet (Fig. 12(d)) using Cu interlayer. As depicted in Fig. 12(e), the through cracks (0.5 mm) were generated as the Ti-Fe-Cu intermixing (Fig. 12(f)) was not prevented due to the lower melting point of Cu when compared with Fe. Hence, Fe-Ti and Ti-Cu based IMCs were formed. The welded joint presented a 27% decrease in UTS and a 23% decrease in the impact energy as compared to the parent metal whereas the fracture surface was uneven with intergranular morphology. However, the welded joints showed a significant decline in the bending fracture load wherein the steel side toe appeared to be the weakest part (Fig. 12(g)). A significant enhancement in the UTS (627 MPa) of TC4/SS301L joint was achieved when Zhang and his colleagues [171] employed multiple interlayers of Ta/V/Fe materials that prevented the Ti-Fe intermixing. The two-pass laser beam was focussed on the Ta and Fe layers that prevented the complete melting of the V interlayer, that eventually promoted the strength. As per research work [172], Ti and Ta form BCC solid solution and similarly Ti and V are also completely miscible [173]. The FZ at the Fe-V interface showed the presence of γ-Fe+(Fe, V) solid solution which was uniform (Fig. 12(h)) and crack free while preventing the formation of brittle σFe phase.

4.2.3. Multiple interlayers In one instance, the use of multiple interlayers that have good compatibility with Ti-SS combinations form the basis for a viable solution to prevent IMC formation and enhance the joint strength to be equivalent to the UTS of the interlayers [165]. Multi-pass-narrow-gap welding technique made up of multiple interlayers of Ti (TA1), Cu (HS201) and Fe (ER50-6) fillers were used to investigate their corresponding influence on the transition zone between butt-welded CP-Ti/ Q235B bimetallic sheets that are commonly employed in sealing pressure-bearing welded structures [33]. When the process was compared with GTAW [56], the area of transition zone was reduced considerably, the size of FZ was 1.5–2 times lesser, resulting in a lesser volume of

4.2.4. Hybrid welding Laser welding, when used together with explosively welded multiple interlayers can deliver favourable results. The interlayer materials such as Ta and Nb are very stable and do not form IMC at interfaces between Ti-Nb, Cu-Fe and Ti-Ta. Cherepanov et al. [49] employed CO2 laser welding to join AISI321 and VT1-0 with the composite insert made of Ti-Nb-Cu-SS layer obtained by explosive welding as presented in the optical image of Fig. 12(i). As the formation of IMC was completely avoided, the highest joint strength of 476 MPa was observed, so signifying the efficiency of the hybrid process. When the authors replaced Nb with Ta, a reduction in UTS value of 417 MPa was observed [174]. In other work, laser cold metal transfer arc hybrid-welding was carried 16

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requires a certain degree of mutual solid solubility to promote the feasibility of the joint. According to the Ti-Al phase diagram as depicted in Fig. 14(a) [179], the solubility of Al in Ti is 13% while the solubility of Ti in Al is close to 0% at 500 °C. The TiAl3 phase forms at the Al-rich side at 2% of Ti. There is a negligible possibility that a certain amount of Al could be present in Ti, whilst not forming the IMC. The control of such minor composition during welding-brazing is extremely difficult and efforts are underway to limit the formation of many brittle IMCs such as Ti3Al, TiAl, TiAl2 and TiAl3. Various techniques have been employed to reduce the detrimental IMC formation such as laser offsetting towards Al or Ti side (Fig. 14(b)), using either butt joint or lap joint configuration (Fig. 14(b, c)), or employing fillers materials and sectioning the end joints as V grooves or U-slots (Fig. 14(d–f)). The effects of these modifications on the UTS are outlined in Fig. 15. In Table 7 we have categorised the fracture and microstructural characteristics of Al-Ti joints based on the location of the offset. Tomashchuk et al. [180] in general showed the effect of laser beam position on the interface morphology. Offsetting the laser beam towards the Ti alloy was found to produce a thick interface, composed mainly of TiAl3, and large defects (Fig. 14(h)) [180]. These defects are formed as a result of capillary trapping and Ti-rich liquid ejection from the melted zone. Furthermore, focusing the beam at the centre of the joint resulted in a reduction of the weld thickness of about 20%, thus affecting the joint strength (Fig. 14(i)). On the other hand, laser offset towards the Al alloy produced a good joint, with a thickness reduction of ≤ 10% and minimal interface thickness (Fig. 14(g)), between 5.4 and 18.6 µm. The hardness of Ti increases due to the heat-treatment which is accompanied by tempering of martensite and dissolution of retained beta phase [181]. In the HAZ, the hardness value decreases due to the lesser amount of martensite available. Laser welding caused an increase

out using a Cu3Si interlayer, resulting in an increase in UTS with an augmentation in the heat input [163]. The composition of the weld and the temperature in hybrid welding are dependent on the melt flow emanating due to the synergistic laser-arc effect. The laser was focussed on the corner of the V groove towards the SS side, resulting in initial and faster melting of the SS side. Furthermore, the liquid convection from the top to the bottom was derived from the arc pressure and surface tension. Whereas, the buoyancy effect caused an upward flow. For a joint with low heat input as shown in Fig. 13(b), a ternary system of Cu-Fe-Si was formed and the melt pool was thinner comprising of a higher volume fraction of Cu3Si. With further rise in the heat input (Fig. 13(d)), a complete mixing occurs and a homogeneous joint with the Cu-Fe-Si-Ti quaternary system is obtained, whereby the FZ comprises of α-Cu matrix made up of Fe67xSixTi33 dendrites. The river-like fracture morphology occurred at the Ti/Cu interface wherein the hardest Cu-Ti2 IMC phase was formed.

4.3. Titanium-aluminium joints The potential application of Ti/Al can be realised for instance in wings made of Ti alloy in which the Ti-alloy crust and Al-alloy honeycomb are welded together [175]. However, direct laser welding of Ti to Al in keyhole mode leads to cold cracking [176]. In laser welding of Ti to Al, the characteristics; size, distribution, morphology and thickness of the undesired IMC phase formed at the Ti-Al interface is governed by the diffusion coefficient, laser welding linear energy, the laser offset location and distance from interface, the choice of weld filler material and the features of the grooves [29]. The diffusion coefficient of Ti in Al is 2.15 × 10−8 m2/s [177] and can sustain at a modest temperature of 600 °C [178]. Fusion welding of dissimilar metals

Fig. 14. (a) Binary Ti-Al phase diagram [179], (b) schematic representation of the laser offset welding towards the Ti side showing different zones [178], (c) lap joint configuration with Al on top [35], (e) the V groove created at 45° for Al and Ti with the usage of filler wire [176] and butt-weld configuration by split-beam laser welding with U slot in the Al side [90]. SEM image and X-ray Al-k map for (g) thin diffusive interface (5 kW; 6.6 m/min; 0.2 mm Al offset), (h) fractured diffusive (5 kW; 6.6 m/min; 0.2 mm Ti offset), (i) malaxated diffusive (5 kW; 8 m/min; 0 centred) [180]. 17

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Fig. 15. The variation of the tensile strength based on the offset location and welding parameters between different combinations of Al-Ti system [43,57,97,142,175,176,178,180,183–185].

joint strength of 80 MPa [188]. The thickness of the IMC layer increases with the amount of incident linear energy. Correspondingly, the higher thickness of the Al-Ti IMC layer increases the chances of crack formation and propagation that could reduce the percentage elongation, yield strength and ultimate tensile strength. With a slight unnoticeable reduction of the UTS, Leo et al. [181] showed that laser offset welding towards the Ti side enhances the elongation due to martensitic tempering and grain coarsening effect after post weld heat treatment at 350 °C. The PWHT at 450 °C causes the diffusion of Ti/Al to increase the quantity of Al3Ti to produce a brittle fracture. The relationship between the linear energy with the shape of the IMC interface and UTS was investigated by Casalino et al. [183] for AA5754 and T40 alloys as presented in Fig. 16. Later on, about 90% joint efficiency was achieved by Casalino et al. [178] for fibre laser welded AA5754 and Ti6Al4V butt joints with a relatively higher offset value of 0.75 mm so as to prevent Ti melting at the interface. Ultra-high linear energy (70 J/mm) caused cracks, geometrical defects [183] and created extended width of FZ, while a decrease in the linear energy to a value of 35.30 J/mm disclosed a rising trend for the UTS due to the formation of a homogeneous and thin, 1 µm, IMC layer.

in the hardness on the AA5754 side due to solid solution strengthening and fine solidification structure accompanied by the dissolution of Mg [182]. However, after performing the PWHT the hardness decreased due to grain coarsening and nucleation above the Al solvus line [181]. At an offset of 0.2 mm toward the Al side, Nikulina et al. [29] showed that the hardness at the interface was proportional to the laser power which governed the thickness of the IMC layer (200 µm). The hardness of the Ti3Al IMC layer was found to be 490 Hv that causes brittleness in the joint, but can be avoided by limiting the amount of heat supplied and restricting welding speed. Casalino et al. [178] showed that laser offsetting towards the Ti side causes slight precipitation hardening in the FZ of Al side due to refinement in grain size after the precipitation of Mg at AA5754 grain boundaries. Furthermore, the FZ in the Ti side also registered higher hardness due to the rapid cooling effect causing the formation of acicular martensitic structure (α′). Together with experimental investigations and inspections, the modelling and numerical simulations are equally important to predict and comprehend the temperature distribution, weld geometry, diffusion, IMC growth and welding performance [188]. With regard to numerical simulations, Dal et al. [35] employed multi-physics simulation based on heat transfer, fluid flow and mass transfer to validate the thickness of the IMC layer experimentally with a margin of 25% error. The margin of error was attributed to the assumptions made on input parameters, diffusion coefficient and activation energy while the aspects of the IMC grain growth parameters were neglected.

4.3.2. Laser offset welding Al-side An enhanced joint strength was validated by Sahul et al. [184] as they employed disk laser at an offset of 300 µm towards the AA5083 side to obtain a UTS of 170 MPa without employing any grooves or fillers. In other work, a 0.2 mm offset at the 1424 Al side with minimum linear energy produces a very thin, 1 µm, TiAl IMC layer [29]. A closer look at the interface through transmission electron microscopy reveals 3 regions comprising of a continuous TiAl IMC layer (Fig. 17(a, region 1)) on the boundary of VT6S alloy, Al3Ti and liquid Al phase. The TiAl phase forms when the liquid phase and alpha phase Ti interact and are supersaturated with Al. A separate region towards the Al side containing Al3Ti is formed (Fig. 17(a, region 2) from TiAl as it interacts

4.3.1. Laser offset welding Ti-side Laser welding with offsetting towards the Ti side can be advantageous as the higher reflectivity of Al reduces the process efficiency, reactivity and low melting-point, causing spattering. Hence it generates lesser spatters than Al offsets. Numerical simulations that were validated experimentally revealed that an offset distance of 250 µm (Tiside) for AA5754/Ti-grade-2 at a welding speed of 10 mm/s produces a 18

Filler

Al-12Si

AA5754:T40 [176]

19

2, 1.8 3

– Al-12Si

Al-12Si

U Slot

AA6056:Ti6Al4V [90] 5A06:Ti–6Al–4V [186]

5A06Al:Ti–6Al–4V [57]

AA6056:Ti6Al4V [187]



1.5

Al-12Si

5A06:Ti–6Al–4V [185]

Laser lap welding 5052:Ti-6Al-4V [175]

2



AA5754/Ti6Al4V [180]

2, 1

2, 1.8

2.5, 2, 2

2, 1.2, 2

1

5087



4, 3

AA5083:Ti-Grade-2 [184]

AA2024:Ti6Al4V [142]

Offset Towards Al side AA1424:VT6S [29] –

1.5



AA5754:T40 [183]

3, 1.6, 3

2



3, 2

Thickness (mm)

AA5754:Ti6Al4V [178]

Offset Towards Ti side AA5754:Ti6Al4V – [43,182]

Materials

Multimode CO2, CW

Nd:YAG

CO2

Nd:YAG CO2

Nd: YAG

Yb:YAG

Disk laser, CW

Nd:YAG, PW

CW

Yb:YAG, CW

Yb:YAG, CW

Yb- Fibre CW

Yb- Fibre, CW

Laser Type

Overlap, Ti Side

Butt

Butt

U slot, Butt Butt

Butt

Butt

Butt

Butt

Butt

Butt

Butt

Butt

Butt

Joint Configuration

2000







180 kJ/m

3500 2400

2400

5000

2000

25 J

4400

3600

50 J/mm

1200, 35.30 J/ mm

1200

Optimum Power (W)/Others

0.4 mm Al side

– 0.8 mm, Al

0.4 mm Al side

0.2 mm, Al

300 lm, Al side

Al side, 0.3 mm

Al side, 0.2 mm

Ti side, 0.9

0.75 mm, Ti

0.75 mm, Ti

Ti side, 1 mm

Offset Location

1.25 m/min, TEM01



Rectangular spot, V-shaped groove with a 45°

0.22 m/min 0.35 m/min groove, 55° angle Al

0.5 m/min, 45° angle V.

6.6 m/min

30 mm/s., focused at +2 mm

200 to 240 mm/ min, Ar

70 mm/s

0.55 m/min 45 ° V groove at Ti and Al side

defocusing (−2 mm)

2 m/min

2000, 0.4 mm spot, Ar & He,

Welding Parameters

Rod-like IMC compounds distribute in Al weld,

Stable fatigue propagation, brittle IMC was observed at interface.

– Temperature gradient cause uneven distribution of reaction layer morphology Four different zones: parent metal, bond zone, columnar crystal zone, equiaxed crystal zone

Increase of linear energy at Al side increases the quantity of Al dissolved in Ti-rich melted zone High heat input causes excess IMC and a mass of porosities

Al0.95Mg0.05, Al3Ti, Ti3Al, Al2Ti, α-Ti

Thin continuous layer adjacent to the Ti alloy with TiAl3 (thickness of 1 – 6 mm) Well refined spherical grains in FZ

FZ was martensitic in the Ti & fine dendritic in the Al side. HAZ partially martensitic for Ti HAZ (equiaxed primary α and inter-granular β phases) FZ (equiaxed α, retained β, and acicular martensite α1) Linear IMC with some Ti lamellae protruded from the interface

Microstructural Characteristics

Smoot brittle and rougher ductile

4 fracture types in partial interface, seam, interface and in the porosity. Transcrystalline Serrated and clubshaped interfacial morphology suppresses crack At higher heat input fracture at porosities in FZ or at the interfacial reaction layer Brittle, Interface

Microvoids caused stress concentrators, Brittle intergranular and transgranular Cleavage fracture mode





Transgranular and intergranular fracture

Interface linear profile better properties than curvilinear

Brittle larger and undeformed flat cleavage areas

Fracture Characteristics

Table 7 The microstructural features and fracture characteristics for dissimilar Ti-Al based laser joints classified according to the process modification employed.

Ti WZ

5–20



0.05 to 0.4

– 1–10

(continued on next page)

TiAl and TiAl3, Ti3Al

Low heat input, TiAl3 and at high heat input TiAl3, TiAl, Ti5Si3, and Ti3Al –

FZ



– –



– Broken in the seam

FZ IMC reaction layer

5.4–18.6

432–378

TiAl + TiAl2, TiAl and Al3Ti, Ti3Al

Al3Ti continuous thin layer





Al-rich melted zone

Near AA2024 side. Ti interface

1000

2–25

1

1–2

50

IMC Thickness (μm)

TiAl3 & TiAl

Si rich, Ti5Si3 and T2, Ti(Al,Si)3

T40/MZ interface



TiAl3

Continuous TiAl, TiAl3, TiAl2, and Ti3Al



IMC layer

Ti side (Ti3Al and TiAl) Al side Ti5Al11 and TiAl2, Al3Ti

IMC Phase

Interface

Fracture Location

M.M. Quazi, et al.

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1 TiAl2

4.3.4. Laser hybrid welding Laser arc hybrid welding can transfer the heat input to the Al side more efficiently because Al does not absorb the laser energy efficiently [190]. Laser cold metal arc transfer (CMT) hybrid welding process was utilised by Gao et al. [97] to obtain a base metal (BM) joint efficiency of 95.5%. The joining process was welding-brazing wherein the Al side underwent welding and the Ti side was brazed with a liquid pool, known as atomic diffusion. It was revealed that the formation of the IMC layer (0.7 µm) at a laser power of 2.5 kW as displayed in Fig. 18(a) was sufficient to form a metallurgical bond in the window of certain process parameters that contribute to specific heat input in the range of 82–98 J/mm as determined in Fig. 18(b). As revealed schematically in Fig. 18(c), when sufficient heat input is supplied, the molten pool flows outwards and upwards to fully cover the Ti interface to achieve a sufficient reaction. The Ti atoms dissolve in the molten pool and subsequently react to form TiAl2 as the cooling rate is fast enough to suppress the detrimental TiAl3 formation. Inadequate heat input causes root defects while excessive laser power causes heat accumulation that lower the solidification rate in the top corner towards the Ti side (Fig. 18(e)) causing it to melt and allowing Ti atoms to move further in greater concentration resulting in the formation of thicker continuous TiAl2 layer and TiAl3 that follows L + TiAl2 → TiAl3. Apart from hybrid

2.5 m/min 20° angle bevel edge at Al side Ti I groove 2500, 83–98 J·mm−1 6061-T6:Al-12Si: Ti6Al-4V [97]

Al-12Si

2.0

Fibre, CW

Butt

Laser focus at the edge of Al bevel

Weld area was annular with no reaction due to the absence of fusion FZ composed of equiaxed dendrites consisting of αAl matrix and a small amount of precipitates (Al12Si eutectic) in grain boundaries 2 and 4 mm laser spot sizes 3.1 J – Laser hybrid welding AA1100:Grade 2-Ti [109]

25–250 um, 75 um

Nd:YAG, PW

Lap

Nill‘

supersaturated solid solution of Ti

H2 pores, tearing ridges and a few dimples, brittle feature

Fracture Location Fracture Characteristics Microstructural Characteristics Welding Parameters Optimum Power (W)/Others Offset Location Joint Configuration Laser Type Thickness (mm) Filler Materials

Table 7 (continued)

4.3.3. Using lap joint, pre-cut grooves, split beam and filler metal For a strong metallurgical contact, a controlled obvious reaction layer is the key. By utilising a 55° groove angle at the Al side, Chen et al [186] studied the effect of interfacial reaction layer morphology on the crack susceptibility and UTS. It was demonstrated that crack initiation sites have a higher possibility of propagation at the Al side because Ti has higher fracture strength. The reaction layer is quite thin and crack propagation herein will lead to plastic deformation at Ti and Al causing a significant increase in deformation energy. Hence, the fracture characteristics are dependent on the Al-reaction layer interface morphology as they originate in the seam of the Al side. They demonstrated that the UTS is higher while the crack propagation is lower for cellular/serration shaped (Fig. 17(i–l)) followed by lamella-shaped but overall it is better than thick or no reaction layer. In order to improve the homogeneity of the reaction layer in a continuation of their work, Chen and peers [185] employed a rectangular shaped pulse with a 45° V shaped groove on Al and Ti ends to obtain an interface parallel to the isotherms of the temperature field, which resulted in a UTS of 278 MPa, higher than their previous study. In order to reduce the detrimental interfacial effect of the IMC layer, laser lap joint configuration was employed between 5052 and Ti-6Al-4V to obtain a UTS of 184 MPa by increasing laser power and decreasing welding speed to provide wider joint with sufficient amount of interfacial reactions [175]. Too high or too low scanning speed and power lead to failure in the interfacial reaction layer while the optimum parameters lead to fracture in the Ti WZ, in the form of brittle fracture in the Ti side and ductile-shear fracture in the Al side. Vaidya et al. [187] employed split beam laser to melt the AA6056 side with u-slots to obtain a brazed joint with Ti6Al4V. The fatigue crack growth performance was lowest at the interface, and a parallel crack that hits the interface (90°) experienced a change of direction along the interface causing immediate failure. Experiments have indicated that the angle (β) at which the filler is fed to the workpiece should be kept at 25-45°. Tomashchuk et al. [176] obtained a UTS of 200 MPa with a joint efficiency of 90% for double half-spot-welded joints AA5754 and T40 by employing Al-Si filler along with 45° V groove for both Al and Ti at an offset of 0.9 mm towards Ti. Fig. 17(e-f) revealed that eutectic structure of 4047 accumulated at the interface and formed 1.8–2 µm Si-rich layer with petal-like structure (TiAl3 + Ti5 Si3).

Failed through Al flyer Weld metal, FZ



IMC Phase



IMC Thickness (μm)

with the liquid Al phase. The UTS as obtained by friction stir welding between AA2024 and Ti6Al4V had a joint efficiency of 62% [189]. Laser welding was able to enhance the joint strength to about 290 MPa.

20

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Fig. 16. Evolution of the IMC interface and relationship of tensile strength with the linear energy.

mechanism entirely prevents the formation of IMC as the joint is dependent on the induced plastic deformation. Wang et al. [109] obtained a weld strength greater than that of Al BM between AA1100 and grade 2 Ti by employing a peel test. They demonstrated that the higher spot

welding, a new welding method known as laser impact welding has been introduced, that can join very thin sheets (pacemakers, batteries) and foils together by using a pulsed laser to generate confined plasma (1000 m/s) to throw the thin foil towards the target sheet. The welding

Fig.17. Bright-field electron microscopy image of the zone of interaction (a) of welded VT6S and 1424 alloys and electron microdiffraction patterns of regions 1 – 3 (b – d), respectively [29]. T40/MZ interface for optimized sample interface morphology (e) and compositional maps of elements Al (f), Si (g) and Ti (h) with an illustration of EDS signal variation across the interface [176]. Fracture surface and interface microstructure for (i) serrated/cellular shaped reaction layer showing crack deflection and (j) second crack tip stopped propagation of reaction/seam interface, (k) club-shaped reaction layer, (l) lamella-shaped reaction layer comprising of many tear ridges signifying cohesion between seam and reaction layer [186]. 21

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Fig. 18. Tensile strength as a function of (a) laser power and (b) heat input. Schematic drawing of the mechanism for IMC layer growth for joints with laser power (c) 2.5 kW, (d) 1.5 kW and (e) 3.0 kW [97]

size can increase the weld areas, causing lesser damage to Al flyer due to lesser power density while the Ti side experienced an increase in hardness due to twinning induced plastic deformation. Higher impact velocity resulted in more waves in the microstructure with a shorter wavelength and smaller spot size, which contributed to higher amplitude. Laser-based hybrid and impact welding have disclosed encouraging results and adequate joint strength to substantiate their usage for exploring further applications.

maximum joint efficiency of 85.1%. As the laser offset distance decreases, the joint strength also decreases because the intensity is too strong to melt Ti that intermixes with Mg to subsequently vaporise it and becomes visible in the form of spatters on the weldments [192]. If a filler with a lower Al content is utilised, then according to Fick’s law of diffusion Al atoms tend to migrate from the Ti side to the Mg side and as the temperature is lowered below 437 °C then eutectic Mg17A12 form in α-Mg. As described schematically in Fig. 19(c), as the offset distance is larger than 0.4 mm from the Mg side, the amount of temperature available at the interface is not enough to instigate the diffusion and melt the Ti side [193]. The fluid flow is restricted by the solid Ti causing unstable vortex that destroys the uniformity of the weld. As the offset is reduced below 0.4 mm, the power density is enough to instigate a pool flow due to the combined effects of recoil forces, gravity and buoyancy that cause curving of the Ti-weld interface (Fig. 19(b) [193]. Hence, laser offset at 0.3 mm encourages the Mg evaporation and the enhanced Ti-Mg intermixing to initiate the eutectic reaction.

4.4. Titanium-magnesium joints In the pursuance of exploring the engineering applications of Ti/Mg joints, researchers have now initiated efforts in investigating the effect of various hybrid welding techniques and the use of fillers to produce joints of acceptable quality. The main challenge is the substantial difference in thermos-physical properties of Ti and Mg at which Mg evaporates at 1091 °C and can render laser fusion welding inapplicable. Additionally, the binary Ti-Mg phase diagram (Fig. 19(a)) shows that Ti and Mg are immiscible whereby no reaction or atomic diffusion is possible after solidification [191]. Acceptable Ti/Mg joint strength can be obtained by employing either laser offset welding or laser welding brazing hybrid process while their corresponding features and mechanical properties are summarised in Table 8. The details and explanatory notes are provided in the following sections.

4.4.2. Laser lap welding brazing Laser welding-brazing is an emerging process for joining immiscible materials that were inspired to join dissimilar materials that are struggling to make their mark in real-world applications because of joints’ poor mechanical strength. Herein, welding is performed for low melting point materials in lap joint configuration, while brazing is implemented in the high melting point materials. Direct fusion welding of Mg/Ti produces weak bonds with poor joint efficiency because Mg and Ti are immiscible and do not form any interfacial or diffusion layer.

4.4.1. Laser butt offset welding brazing Laser offset welding brazing of AZ31B/Ti6Al4V has stipulated a 22

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Fig. 19. (a) Binary phase diagram of Ti-Mg. Bead formation mechanism for the offset distance (a) less than 0.4 mm, and (b) greater than 0.4 mm [193].

Hence, filler elements that possess intermediate solid solubility with both Mg and Ti are required. Yet the interfacial layer made of IMC is required to be less than 10 um as it could be favourable for the mechanical behaviour. The formation of diffusion reaction layer can enhance the UTS of joints by transforming the mechanical bonding into metallurgical one as well as by preventing crack propagation [195]. For instance, the employment of AZ91 filler (9 wt% Al) for AZ31B/Ti6Al4V joint enhances the joint efficiency of AZ31B filler (3 wt% Al) from 22.9% to 47% with brazing filler [195]. This is because the bonding mechanism in direct fusion welding is mechanical interlocking (Fig. 20(a)) while laser welding-brazing creates a 1 μm thick reaction layer (Fig. 20(b)) at the Ti side leading to a metallurgical bond. Ti3Al reaction layer was formed and the diffusion controlled Ti-Al and restricted due to rapid thermal gradients. A positive Ti-Mg molar enthalpy indicated that no mutual reaction occurred (Fig. 20(c)), exhibiting that in an Mg-Ti-Al ternary system the Al has the tendency to diffuse into the zone with lower Al and higher Ti contents. As determined in Fig. 20(f), the increase in laser power at lower welding speed augmented the spread-ability and wetting of filler that led to enhanced atomic diffusion. A further increase in laser power deteriorates the joint strength because higher heat input evaporates the Mg filler resulting in weak interfacial bond [191]. To further enhance and control the interfacial reaction, Ni was selected as an interlayer (1.9 μm) with AZ92 (8.3–9.7 wt%) filler [196]. Ni-interlayer enhanced the spread-ability and wetting behaviour of the filler leading to stabilisation in the welding process. The addition of Ni interlayer enhanced the tensile shear from 2057 N (AZ91) to 2387 N (AZ92). The thickness of the reaction layer was found to increase from 2.08 to 3.22 μm with an increase in laser power from 1100 W to 1700 W. The microstructural evolution at different zones was schematically described in Fig. 20(g–l). Upon direct laser irradiation, the Mg filler melted (Fig. 20(g)) while the Ni coating dissolved and diffused in the molten Mg (Fig. 20(h)). Wherein, the laser heat input was not adequate to melt the Ni layer, an intermediate zone was formed since the liquid flow was not profound. At the direct radiation zone, the Al element of the filler diffused to the Ti side to form Ti3Al precipitate upon cooling (1180 °C) as disclosed in Fig. 20(j). In the intermediate zone, the Ni and Al atoms were both in the liquid state and dissolved within each other (Fig. 20(i)). With a further reduction in the temperature below 650 °C, the liquid AZ92 began to solidify and induced a reaction of Mg with Al and Ni leading to the formation of Mg-Al-Ni ternary compound near to the interface intermediate zone while Al-Ni phase formed at the interface. The Mg-AlNi ternary compound thereafter grew from dendrite to needle-like structure with an increase in laser power as revealed in Fig. 20(l) [196]. The reaction layer thickness increases with a rise in interface temperature and diffusion time [197]. Hence, the offset distance has a

significant effect on the reliability of Ti/Mg joints as it controls the resulting temperature and diffusion at the interface. The hardness of the interface is essentially dependent on the amount and random distribution of the IMCs that varied with the variation of power. 4.5. Titanium-nickel joints Nickel and its alloys are widely employed in the high-temperature aerospace applications where oxidation resistance is of utmost importance. Few investigations have paved the way for Ni-Ti dissimilar welds. In earlier work carried out in 1976 by Seretsky and Ryba [198], spot welding of Ti to Ni revealed cracks and incomplete mixing of molten metals. This was seconded in later work by Chatterjee et al [199,200], as they discovered micro-cracks, brittle IMCs of Ti2Ni and TiNi2 with macro-segregation. However, until recently only, Chen et al. [201] employed fibre laser welding technology at higher power and a higher welding speed to obtain crack free welds between butt welded Ti-6Al-4V and Inconel 718 alloy. The laser beam, when offset towards the Inconel side, caused less vigorous convection flow in the molten pool and a significant reduction in the melt area of the Ti side as noted. The decrease in the intensity of the Marangoni convection currents led to lesser mixing that in return alleviated the IMC formation. Additionally, the higher thermal conductivity of Ni meant that the heat could dissipate faster resulting in a wider FZ in conjunction with lower thermal gradients. To highlight the feasibility of employing low power fibre laser for welding T-joints, Janasekaran et al. [54] employed a 50% overlapping factor to obtain a maximum fracture force of 150 N for Ti6Al-4V - Inconel 600 joint. The hardness of the joint in FZ was perceived to intensify with overlapping ratio and was significantly higher than that of the BM due to crystallographic mismatch and formation of brittle NiTi and NiTi2 IMCs. It was established that overlapping was the most influencing factor affecting fracture force followed by welding speed and laser power. 4.6. Titanium-niobium joints Laser welding of niobium (Nb) metal is only being discussed in a few reports lately. Recent research has been carried out because of their state-of-the-art application in high-performance structures for particle accelerators [202]. Niobium is considered as a refractory material having superconducting properties. Nb is unreactive and does not form IMC compounds in the Ti-Nb phase diagram (Fig. 21(d)). Torkamany et al. [38] studied the keyhole and conduction characteristics of dissimilar welding of Ti-6Al-4V and niobium through direct visualisation of the laser using high-speed imaging. Because of the difference in thermal conductivities of materials, the Ti (6.7 W/m K) sustained the 23

24

Ti–6Al–4V:AZ31B/2 [192]

AZ31B (1 mm)

Laser butt Offset Welding-brazing Ti-6Al-4V/AZ31B-T5/2, 3 No filler [193]

AZ31, AZ91 filler wire

Ti–6Al–4V:AZ31B/1.5 [195]

Fibre, Continuous, Butt joint

Fibre, Continuous, Butt Joint

Fibre, Continuous, Lap (Mg on top)

Fibre, Continuous, Lap (Mg on top)

AZ91 filler wire (1.2 mm)

Ti–6Al–4V:AZ31B/1.5 [191]

Laser Type/ Configuration

Fibre, Continuous, Lap (Mg on top)

Filler/ Interlayer

AZ92 filler (1.6 mm)

Laser lap Welding-Brazing Ni-coated Ti–6Al–4V:AZ31B/1 [194]

Materials/Thickness (mm)

2.5 kW, 2.5 m/ min. offset 0.3 mm Mg side 2 kW, Offset 0.2 mm Mg side, 2.4 m/min



2.4 kW, 0.5 m/ min

1.3 kW, 20 m/ min

Optimum Welding Parameters

Not reported

Not reported

α -Mg and Mg–Al eutectic (Mg17Al12)

less than 1

less than 1

2.08–3.22

IMC Thickness (μm)

Some lamellar and granular mixtures with a composition of α -Ti, α -Mg, and Mg17Al12

AZ91 - (α-Mg + β-Mg17Al12) + Ti3Al

Serrate-shaped formed, Ti3Al

Ti3Al phase, Al-Ni phase, Mg-Al-Ni ternary compound

Microstructure/IMC

Smoother area for low UTS, scraggy, small convexes, with smooth areas for higher UTS Scraggly appearance

AZ31-smooth with adequate plastic deformation AZ91serrate-shaped

Cracking with small particles attached (Ti3Al)

Crack extended along the interface at the seam root zone with dimples

Fracture Characteristics

Table 8 A summary of mechanical properties, microstructure and fracture characteristics of laser offset welded and laser brazed Ti-Mg joints.

Ti/FZ interfacial layer

Ti/FZ interfacial layer

Interface Ti side

Interface & FZ both

Interface & FZ both

Fracture Location

When the offset increases, the joint strength decreases

Larger offset prevents reaction and diffusion layer formation

Fracture transformed form interface failure to FZ failure at 1300 W, Reaction layer thickness increases with increasing Power At the high welding speed of 1 m/min, the reaction layer was the thinnest due to the lowest heat input Increase in Al content in filler enhanced the diffusion

Parameters Relation

Tensile 200.3 MPa, Joint Efficiency 85.1%

Tensile 266 MPa

Tensile shear Fracture load 2075 N, Interfacial Hardness 62 Hv, 50% Joint efficiency Tensile Shear Fracture load 1050 N @ AZ31 filler, 1947 N @ AZ91 filler, Interface Hardness 275 Hv, Joint Efficiency 47.3%

Tensile shear Fracture load 2387 N, Interfacial Hardness 82 Hv, 88.5% Joint Efficiency

Mechanical Properties

M.M. Quazi, et al.

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Fig. 20. SEM images and the corresponding EDX scans for (a, c) AZ31B filler, (b, d) Al-rich AZ91 [191] (e) molar potential with decreasing Al content [195]. (f) Tensile-shear strength profile of AZ31B/Ti6Al4V samples with increasing laser power [191]. Schematic of joining mechanism: (g), (h) melting of filler and Ni coating, (i) dissolution and diffusion of Al atoms, Ni atoms at the intermediate zone, and Al atoms, Ti atoms at the direct irradiation zone, (j)–(l) solidification of interfacial zone at different temperature ranges [196].

rapidly expand in the confined transparent layer, thus creating a high surface pressure. As a result, shockwaves are generated, which propagate in the hump to deform and produce a solid-state bonding due to plastic shear deformation under great impact pressure (Fig. 22(b)). But since the laser Gaussian energy is higher in the middle, greater impact was witnessed in the centre of the bulge. The lump due to impact pressure contacts the base plate at 0° but as the collision spreads radially, the collision angle rises and a wave-like structure is formed as shown in Fig. 22(d). Consequently, as the angle increases, spot welding is accomplished with some amount of springback due to the release of the shock pressure (Fig. 22(e)). They reported that the lap shear strength was found to intensify with the increase in pulse energy because the weld contact region enlarges. Direct laser offset fusion welding of T2-Cu with TA18-Ti was carried out by Zhao et al. [208] to report a joint efficiency of 61% by obtaining a UTS of 151 MPa at an offset distance of 0.45 mm towards the Cu side. The WZ contained Cu solid solution and a brittle IMC layer of 25 μm containing TiCu, TiCu2, Ti2Cu and Ti3Cu4, and the failure occurred in a quasi-cleavage mode. Using Nd: YAG laser, Zhao et al. [206] reported that the welding zone of Ti/Pb joint was composed of FZ in the upper part and brazed zone in the lower part as presented in Fig. 22(i). The FZ was comprised of Ti4Pb IMC formed due to eutectoid decomposition and peritectic reaction. The Ti element in the FZ was unable to transport to the Pb zone due to the solubility gap of Ti in liquid Pb. Chan et al. [27] investigated the LTW of CP-Ti with Poly (Ethylene Terephthalate) PET and quantified the joint strength based on the contact area as determined by ImageJ analysis software. A maximum shear strength of 3.2 MPa was obtained. LTW prevents the risks associated with adhesives such as leaching extractable products, long-term

laser energy for a longer duration while Nb (53.7 W/m K) dissipated the energy to the bulk metal quite faster as seen in the top and cross-section views of the single spot laser pulse (Fig. 21(b, c)). In other work, Torkamany et al. [203] determined the optimum parameters (power; 1.5 kW, pulse duration; 3 ms) of pulsed Nd; YAG laser welding of Ti-6A4V (0.85 mm) with Nb (1 mm) plates. The fracture occurred at the Nb side while exhibiting a UTS equivalent to BM, being 269 MPa. In the FZ, the average hardness declined from harder Ti to softer Nb (Fig. 21(d)). It was also reported that Nb did not react with the alloying elements Al and V. The authors recommended that laser energy be directed more towards the Nb side to gain more uniform penetration. These reports suggest that Nb is an ideal candidate to avoid IMC formation in joints and can effectively serve as an interlayer. 4.7. Welding of Ti with other materials Quite fewer attempts have been made to investigate the laser welding of Ti-alloys with polymers [205], Lead (Pb) [206] and copper (Cu) [207]. Recently, a novel laser impact spot welding (LISW) was utilised to weld thin plates of pure Ti and Cu without melting of the constituent materials [207]. The two plates were joined by the solidstate welding mechanism (Fig. 22(a)) under low temperature, and elevated impact velocity and high angle. The pre-deformed flyer plate is attached to the base (Cu) plate with an adhesive. Nd: YAG laser with a pulse energy of 1800 J/mm and a pulse width of 8 µm impacted the hump while it is transferred to the transparent confinement layer. Thereafter, the laser beam strikes an ablative absorbent layer on the hump and instantaneously vaporises the material. As the temperature increases, plasma is formed then absorbs the incoming laser energy to 25

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Fig. 21. (a) Ti-Nb dissimilar weld cross-section obtained for laser pulse energy of 9 J and duration of 6 ms. The effect of a single pulsed spot on spot intensity on (b) Ti and (c) Nb along with their corresponding penetration depths [38]. Binary phase diagrams of (d) Ti-Nb signifying no existence of any IMCs and (e) the variation of hardness across the welds [204].

5. Difficulties in laser welding of titanium and main metallurgical defects

stability and slower processing. The transmitted laser energy through plastic heats the underlying metal, which then heats up the contacting surface of the plastic, causing it to decompose. Thermal decomposition of the heated plastic surface generates solidus, liquidus and gaseous products at the plastic-metal interface. The gaseous products form bubbles, and the expansion of these bubbles creates a high local pressure, which forces the molten plastic liquid products to the metal surface. After re-solidification, mechanical interlocking, chemical reactions and Van Der Waals’ interaction in the form of stable bonds are formed between the plastic and metal interface. For instance, it is to be noted that laser power followed by scanning speed, material transmissibility, clamping pressure and absorption properties at the interface are essential parameters [209]. Wang et al. [210] investigated the joint performance between Ti and PET made by CW diode laser under lap joint configuration. The tensile strength was inversely proportional to the scanning speed and similar trends were observed for a particular range of laser power. For applications involving bio-encapsulation of medical implant devices, Mian et al. [211] predicted the failure load of diode laser welded Ti-polyamide joint using FEA model and observed that the failure load to be in agreement with the experimental results. Chen et al. [212] used pulsed Nd: YAG laser to join Ti and PET, and concluded that the degradation of PET can be minimised by properly controlling the waveform of the laser, i.e., ramping up/down. According to the Hall-Petch relation, the hardness and strength of metals increases with decreasing grain size. Compared with the grains in the unaffected BM, the grains in the recrystallisation and fine-grain zones were finer. In addition, the finest grain size was found in the fine-grain zone. Fig. 23(a) and (b) depict the UTS and hardness variation of different materials welded with Ti.

Ti and its alloys are extremely sensitive and reactive to gases present in the atmosphere at high temperatures. Albeit, in the open literature argon gas, has been utilised extensively to enhance the efficiency of joining and improve the mechanical performance of Ti and its alloys. Nonetheless, an impurity like O2, N2 and H2 that might exist in small amounts as traces severely impede the soundness of the weld and HAZ [213]. It was reported that traces of 0.005%O2 in the shielding gas composition lead to the formation of rutile TiO2 in the weld joints indicating the extreme sensitivity of Ti as a metal for welding. Additionally, TiN is formed when N2 in the composition exceeds 0.008%. However, the gravest danger lies in hydrogen-based porosity, as with increasing temperature owing to laser irradiation, the solubility of H2 increases. The water molecules and nascent hydrogen present in the atmosphere dissociate when subjected to very high temperatures (2000–3500 K) [214] and follow the dissociation reaction as below: H2O↑ = H2↑+1/2O2↑

(1)

H2↑ = H↑+H↑

(2)

The welding defects are strictly influenced by the processing parameters. Ti and its alloys are not susceptible to solidification-based cracks in the FZ and liquidation cracks in the HAZ due to the absence of the second phase particles, impurities and dispersions that result in a joint without discernible cracks in the WZ. The choice of filler material and pulse energy directly affects the weld quality. For dissimilar joints, the control of the thermal and diffusion processes is essential to avoid structural and microscopic defects. Lack of fusion and penetration is entirely dependent on the applied parameters and the joint gap distance as more loss of energy is conceived with an increase in the joint gap 26

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Fig. 22. (a) The experimental principle of laser impact spot welding (LISW), (b) Optical micrographs of the cross-section view of Ti/Cu joints produced at different laser pulse energies: (b) 565 mJ, (c) 1550 mJ. Schematic illustration for the formation process of the laser impact spot welded joint; (d) before LISW; (e) the hump bulges and collides onto the base plate at a high speed; (f) the hump contacts the base plate first at 0°; (g) welding phase; (h) waveform like structure is formed [207]. (i) Chemical composition of different regions in the weld for Ti-6Al-4V-Pb dissimilar joint [206].

higher velocities of the fluid led to ejection of spatters (Fig. 24(a)). Shear streams were formed due to plume ejection at subsonic speed originating from the position of the keyhole to introduce an upward melt flow. Hence, a spatter was ejected from the keyhole wall that accumulated to contribute to the elongated melt. With increasing the welding speed as shown in Fig. 24(b), the elongated molten metal ratio

[215]. Likewise, if a positive defocus of 1 mm is employed, the laser welding mode causes incomplete penetration [216]. Nakamura and peers [217] investigated the spatter formation mechanism for a bead on a plate using 10 KW laser power with the assistance of a high-speed camera and X-Ray detectors. It was determined that at lower welding speeds, an elongation melt was formed in front of the keyhole, and

Fig. 23. (a) Strength of different laser welded dissimilar Ti-based joints and their corresponding weld zones (b) hardness [54,203,207,210]. 27

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Fig. 24. (a) Schematic melt flows in the molten pool and elongated melt geometry leading to spattering (b) effect of welding speed on occurrence locations and sizes of spatters [217]. If the welding conditions are not correctly optimised and suitable parameters are not taken into consideration, defects are introduced into the joint that severely impede the mechanical performance. The commonly observed defects for Ti-based welds have been classified in Table 9 and elaborated briefly in the following subsections.

opportunities for air, gas or hydrogen contamination in near-beta Ti5553 [226]. The maximum allowable size of porosity has been documented as 0.40 mm specified in BS EN 4678, 0.60 mm in BS EN ISO 13919-2 and 0.66 mm in AWS D17.1. However, these standards also consider different criteria, for instance, equivalent projected pore area, diameter of the largest pore, minimum distance between adjacent pores, the thickness of the plates to fit the allowable size of the porosity or the resulting equivalent pore length per given weld length. Further details concerning the mechanism of porosity formation can be found in Ref. [227].

in front of the keyhole inlet side decreases. The plume ejection is deviated to lessen the accumulation of shear forces that return the elongated melt to the melt pool.

5.1. Porosity Deficiency of laser power and incorrect selection of pulse type and duration can produce incomplete penetration while improper use of shielding gas may create cavities, porosity and holes. The shadowing effect of the keyhole wall humps and the rapid cooling of the keyhole tip prior to the collapse of the keyhole lead to a drop in the vapour pressure that allows for the shielding gas to get entrapped [218]. The resulting detrimental effect on mechanical performance is due to residual stresses, while the internal gaps due to incomplete penetration cause internal failure. The porosity and cavities act as fracture site initiators and are the preferred sites for higher stress concentration [219]. The aggregation of pores and their shape affect the fatigue properties [215]. Degidi et al. [220] carried out fractography of fatigue tested specimen of grade 4 Ti and recognised multi-crack initiation sites acting as stress raisers. While fatigue striation is visible in the cavities, the cracks propagated towards the joint centre causing failure in the welded zone. Yu et al. [221] also showed that the small porosities act as crack initiation sites and significantly affect the fatigue life of the welded Ti6Al-4V, however, the mechanical properties such as UTS and hardness were unaffected. The movement of laser vertically toward the bottom position leads to more porosities, however, they can be reduced by increasing the power at higher speeds [222]. Porosity near the FZ/HAZ interface is formed due to keyhole instability and the evolution of hydrogen at the liquid/solid interface. The hourglass-shaped weld pool produced more porosity than the nail-shaped melt pool owing to the downward sweeping front at the root of the hourglass welds [223]. Lastly, the tendency of porosity formation in partially penetrated weld pool is higher because the escape path is only accessible from the top [138]. Increasing the heat input tends to reduce the amount of porosity in Ti–6Al–4V as solidification rate increases allowing more time for gas to escape while increasing the speed increases the porosity due to keyhole instability and collapse [224]. The increase in welding speed can effectively reduce the width of the FZ and might prevent the downward sweeping solidification front which promotes the entrapment of gases [133,224]. The periodic oscillation of the keyhole throughout the welding process also contributes to promoting porosity [218]. However, numerical analysis has shown that weld porosity produces stress concentration that is much more severe than weld seam geometry [225]. With an augmentation of the joint gap, the induced porosity increases for Ti-4.5Al-2.5Cr-1.2Fe-0.1C alloy [215], while the usage of a larger volume of Ti-6Al-4V filler wire allows for more

5.2. Undercut and underfill Undercut and underfill defects are also termed as concavity and have far more detrimental consequences on weld strength than porosity [228]. For typical aero-engine requiring stringent weld quality and have defined tolerance for undercuts, excess weld metal, excess penetration, incomplete filled groove, root concavity, shrinkage groove, etc., within which the acceptable weld is achievable [229]. Defects such as underfill, undercuts, root concavity and shrinkage that are related to the weld profile and geometry are undesirable when the components are subjected to dynamic loading such as fatigue. These defects increase the stress concentrations and hence lead to fatigue crack initiation and accelerated corrosion. Undercutting in laser welding can be formed at the upper and root surfaces and originate at higher welding speed that allows for a lesser time for BM to flow back to the weld toe and fill the gap [224]. Hence, decreasing the welding speed can suppress weld undercuts and root sagging. An increase in power can also leave a drain-like impression along the length of the weld due to enhanced expulsion of the molten material leading to evaporation [230]. The extent of underfill can also increase due to increasing the power density that increases the vaporisation losses. As the vapor escapes, a recoil force is exerted on the weld pool which exceeds the surface tension at the periphery causing underfilling [228]. Some possible approaches to reduce and eliminate undercuts involve employing defocused beam, smaller modulation amplitude, wire feed addition and subsequent machining [231]. By employing pre-extrusion load during laser welding, the underfill defects can be completely eliminated. Apart from underfill, overfill can occur either due to excess filler wire or due to various factors affecting the motion of the weld pool including the longitudinal metal liquid flow, contraction of plates, insufficient welding speed and excess shielding gas flow, etc. Vertical movement of the laser toward the top produces more undercuts than vertical movement toward the bottom because of unstable weld pool dynamics by the action of the gravity [222]. The underfill defect appears above heat input of 328.5 J/cm for Ti-2Al-1.5Mn and when the rate is higher than 28

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used to tailor joints to avoid potentially catastrophic cracks. [244]. The laser welding brazing approach that requires using fillers and offsetting the laser beam towards one of the parent materials to constitute a thin diffusive layer can also produce defect-free joints. Hence, the enhancement of the fatigue crack resistance is entirely dependent on the properties of the interface or diffusion reaction layer [187].

43%, the fracture transforms from the BM to the FZ zone because the cross-sectional area becomes too small [232]. A previous research pointed out that fatigue cycles to failure improves as the underfill radius increases and can be improved further by removing the weld toe [233]. Another research has modified the underfill depth to be under the specified allowable depth mentioned in the AWS D17-1 standard which provides a provision of 7% of the sheet thickness [64]. Besides, the use of filler wire can compensate for the evaporative loss and can reduce the underfill defects [234].

6. Summary and outlook The contemporary survey of the laser welding of Ti-alloys has manifested into an emerging trend shown by the number of publications, consequently leading to a renewed interest for researchers who are contributing to this particular field of materials processing. Although some of the applications related to the biomedical, aerospace, petrochemical and sport industries have been highlighted when considering the laser welding of Ti alloys, further exploration of newer and state-of-the-art applications is necessary in the current technological era. The revitalised appeal of laser welding has transpired materials engineers to continually investigate the welding of newer and dissimilar material combinations that can serve for further applications. The following conclusions can be drawn from the review of dissimilar laser welding of Ti-alloys;

5.3. Cracking Cracking in the WZ may occur because of specific simultaneous conditions such as weld pool contamination, loss of ductility, formation of sensitive brittle microstructure (Widmanstätten), microstructural gradients and higher levels of residual local stress/strain due to melt pool solidification and expansion coefficient mismatch [235]. Likewise, fatigue cracks originate in the vicinity of the other weld flaws [96]. The fatigue crack growth rates (da/dN) for the welded structures are lower than the BM [236]. Cracks also initiate during loading nucleating from the oxidised drops that act as stress raisers. The presence of underfill and porosity further assists in crack development [237]. Cao et al. [238] reported that cracks originate if the depth of the underfill exceeds the threshold value of 6% of the thickness of the specimen. The multidirectional distribution of the martensite phase that grows when subjected to varying the heat input can induce an overall tortuosity that may hinder crack growth [239] but some authors noted that for Ti64 alloys a fully martensitic microstructure is prone to cracking [240]. Coarsening of the martensite platelets by employing slower cooling rates can divert crack propagation paths [241]. IMC in the WZ can generate spontaneous cracks due to thermal mismatch. For dissimilar materials, cracking is favoured due to the loss of the alloying element (Mg), loss of ductility, residual stresses and variation of thermal expansion coefficients. The offset distance does not contribute to cracking in the weld joint but cracks in dissimilar welding are conceived due to phase and volumetric changes because of phase transformation and intermixing [201]. Cracks also arise in gaps caused by lack of fusion [97]. Similarly, cracking tendency is not correlated with any specific trend of applied parameters while cooling rate may or may not determine the corresponding cracking tendency for Ti-Ni based dissimilar welds [201]. For instance, excessive use of heat input or pulse overlapping factor can create cracks during solidification [242] while reducing the critical engineering strain causes failure in the FZ [243]. The application of filler wire can also affect cracking tendency as one research highlighted that grade 5 Ti-filler produces cracks unlike grade 1 [127]. Crack sensitivity of Ti-autogenous welds can be controlled by making alterations in the process parameters but this does not apply for dissimilar materials welds. Crack formation of dissimilar joints is characterised by the uncontrolled growth and reaction of the IMC layer, the amount of brittle phase and its distribution that ultimately cause volumetric and thermal mismatch [134]. For instance, TiNi [47], FeNb, Mg-Al-Ni [196], FeTi [33] and AlTi [182] based IMCs may cause spontaneous cracking and secondary cracks that emerge at the onset of fracture causing instantaneous brittle failure [134]. Nikulina et al. [29] showed that for Ti-Al dissimilar welds, the increase in laser power causes an increase in the width of an intermediate zone that supports cracking while higher scanning speed reduces cracking tendency. Hence, controlling the thickness of the IMC layer becomes essential in preventing crack formation. A summary of different kinds of defects of different materials formed at specific welding parameters for similar and dissimilar joints are tabulated in Table 9. Post welding heat treatment improves the ductility of the brittle martensitic phase and enhances elongation until failure [138]. Welding of γ-TiAl alloy exhibited crack free joints when the pre-heating at a temperature of 750℃ was conducted [150]. Laser pulse shaping can be

1. The welding efficiency of Ti-Ti joint is quite high due to lesser differences in the thermo-physical properties as well as the absence of the IMC phase. Dual beam welding modes have shown beneficial characteristics in improving the tensile strength, elongation and hardness while post weld heat treatment enhances the fracture characteristics by enhancing ductility. These joints retain 60–70% of the tensile strength at high temperatures. 2. The steel-titanium welds are characterised by brittle Fe-Ti IMC joints that cause instantaneous joint failure. To obtain a suitable joint strength, single interlayers of Nb, Cu, Mg and V elements are employed that impede IMC formation, but for each of these interlayers, either Ti or Fe part promotes brittle IMC. Encouraging results have been envisaged when multiple interlayers of Ti:V:Fe are used either through double pass welding or by hybrid explosion welding route to finally prevent the IMC formation. 3. Optimisation of the laser power and scanning speed, that contributes to the overall heat input to the dissimilar Al-Ti joint, can reduce and control the phase, distribution and size of TiAl IMCs. Laser offsetting towards the Al side has provided almost similar joint strength as compared to offset towards the Ti side while the employment of different shapes, angles of grooves and fillers has shown promising augmentation in the joints’ tensile and fatigue strength. 4. Acceptable Ti/Mg joint strength with joint efficiency exceeding 88% has been obtained by employing laser welding brazing process for butt and lap joint configuration involving Mg-based fillers to prevent the intermixing of Ti and Mg. 5. Preliminary investigations have revealed that materials such as nickel, niobium, copper and polymers can be successfully welded. Niobium does not form IMC and retains its joint strength, while laser transmission welding is employed for polymers with titanium, and laser impact spot welding is used to join thin sheets of copper. 6. Failure in laser welded joints has originated either due to structural discontinuities during tensile loading or due to welding defects such as cracks, flaws, holes and porosity. These discontinuities induce either brittle, ductile or mixed brittle-ductile failure that can be comprehended from fractography. Dissimilar weldments have shown brittle characteristics as the joints fail at the IMC interface while joints characterised by ductile failure are envisaged only in the fusion zone where brittle phases are absent. Defects causing stress concentration such as undercut and underfill, due to spattering and evaporation, may become unavoidable. The usage of gas shielding increases for upper and lower sheets, modification of the joint gap and the use of filler wire can help to alleviate such defects. 29

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Table 9 A list of various kinds of fractures occurring for Ti/Mg, Ti/SS, Ti/Al, Ti/Ti dissimilar and autogenous Ti welds. Defects

Materials/Parameter

Titanium-Titanium Dissimilar Joints Porosity [157] Ti-6Al-4V and BTi6431S/Single & Double pass

Reason

Evidence

Single pass welding showed lesser porosity while in double pass welding porosity was found in the overlap zone

Ti-6Al-4V/CP-TiIncreasing the incidence angle from 15° to 25°

Skin was burnt through and unacceptable underfills the heat flow into the skin increases significantly

Steel-Titanium Joints Cracks in FZ [41]

CP-Ti:SS304 CW & pulsed

By employing high cooling rates to reduce IMC thickness cracking could not be prevented in autogenous welds

FZ centreline cracking [34]

TC4:SUS301L/164 W, Pulse width 10 ms

Continuous distribution of Ti–Fe IMC, cracks under the effect of residual stress generate, Cu interlayer prevented cracks

Concavities [163]

SS316:Ti6Al4V, 2 kW, 3 m/min

Insufficient heat input

Cracking due to offset [30]

SS201:Ti6Al4V, Offset 0 & 0.3

Offset towards Ti side produces cracks that led to spontaneous fracture and hence offset towards steel is recommended

Aluminium-Titanium Joints Underfill/Spatter [175]

AA5052:Ti6Al4V, 1800 W 1 m/min

Increasing the welding speed and decreasing power improves the spattering and reduces underfill

Porosity [29]

VT6S:AA1424, 4.2 kW70mm/sec 0.2 offset towards Al

Raising the laser power instigates porosity formation due to vigorous fluid flow

Micro-voids at the interface [183]

T40:AA5754, 1.75 kW, 1.6 m/min

Higher linear energy 75 J/mm causes micro-void formation

Skin burnt through an insufficient connection [36]

(continued on next page)

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Table 9 (continued) Defects

Materials/Parameter

Reason

Lack of fusion [97]

Hybrid weldingAA6061:Ti6Al4V:Al-12Si, 71J/mm

Low heat input occurs due to lesser applied power

Interfacial cracks [186]

Laser welding bracing Ti–6Al–4V : 5A06 offset 0.8 mm, 2400 W, 0.35 m/min

Increasing the thickness of the continuous reaction layer produces weld embrittlement due to residual stress accumulation

Laser welding-brazingTi–6Al–4V:AZ31B, AZ91 Filler/1600 W

Incomplete fusion welding of BM and molten filler occurred resulting in a void

Evaporation [191]

Laser welding-brazingTi–6Al–4V:AZ31B, AZ91 Filler/3200 W

Heat input was too high to produce spontaneous evaporation

Burning through [196]

Laser welding-brazingTi–6Al–4V:AZ31B 1900 W, AZ92 filler

Higher heat input

Inconel 718/Ti6Al4V, 800 W 100 mm/s, Offset at Ti side

Porosity diameter ~15 to 172 μm, crack length 63 and 663 μm. Crack free welds obtained at higher power. Offset distance did not affect the defect formation

Magnesium-Titanium Joints Voids [191]

Titanium-Inconel Joints Porosities and cracks [201]

Evidence

applicability of Ti-based joints have evaded the interests of researchers. It is anticipated that in the near future, the current success of the laser brazing process will allow scientists to demonstrate the maximum capabilities of the process as a prelude to garner the requirements of the industry. The high temperature tensile strength has been evaluated for Ti-Ti joints and has shown encouraging results. In contrast, the ductility and strength of Ti-based dissimilar metals have not been evaluated yet. However, some preliminary work has shown loss of bending and impact strength for Ti-SS joint. There are not many researches engaged in evaluating and improving the impact strength or creep while research on the corrosion, fatigue and bending characteristics of the joints is

The increase in the welding speed and heat input is advantageous in preventing the porosity formation. Laser dissimilar welding of Ti alloys is far from being a mature and reliable process unlike friction stir welding of Ti alloys. The major portion of the work has been focussed on Ti-steel, Ti-Al, and Ti-Mg based joints. However, the different types of Al, Mg and Fe based alloys and their extruded, cast and hydroformed parts in different shapes and thicknesses will probably be welded in the future. So far, laser welding itself as an individual process has not been able to focus on emerging materials. For instance, critical engineering materials like maraging steel, ceramics, composites and porous materials that can amplify the 31

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extremely limited although these characteristics are quite essential for aerospace applications. The information pertaining to the monotonic and cyclic yield strength, strain hardening exponents, fatigue failure crack propagation rates, failure characteristics, etc., are not readily available. Hence, a comparative evaluation cannot be made to determine the process efficiency of laser welding against other fusion welding techniques. The knowledge of the formation of defects and their corresponding relation with the weld pool oscillation in conjunction with the fluid dynamics involved require more in-depth studies [245]. This means that a significant amount of work is required to be accomplished to be able to predict and optimise the laser welding process to produce consistently quality welds. In general, researchers have focussed on studying the effect of single parameters on the joint strength of the weld, whereas optimisation strategies such as response surface methodology, Taguchi, fuzzy logic, artificial neural networks, genetic algorithm, etc., to gain a mathematical model and find knowledge base with regards to mutual interaction of the parameters is still missing. Mechanical properties such as tensile strength and hardness have been investigated largely, however, essential properties that can lead to failures such as fatigue strength, impact strength, fracture toughness and corrosion can be investigated in future research to further strengthen the applicability. For dissimilar welds, weld interface is a region wherein a sharp property change over occurs. Apart from these limitations, it would be wise to conclude that laser welding of titanium alloys remains a growing field with promising opportunities for future research and development.

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