Effect of deposition variables on solidification behavior and mechanical response of LMD-fabricated Ti-6Al-4V alloy clads

ABSTRACT

Laser Metal Deposition (LMD) offers a versatile pathway for near-net-shape fabrication of high-performance titanium components. In this work, Ti-6Al-4V thin-wall structures were fabricated using orthogonally designed experiments to study coupled effects of laser output intensity (800–1400 W), scanning rate (4–12 mm/s) and feedstock mass flow (6–14 g/min). Optimal deposition window was identified at 1100 W, 10 mm/s, 8 g/min, producing geometrically stable tracks with minimum dilution, free of pores. Microstructural characterization revealed epitaxial β-columnar growth at fusion boundary transitioning to refined α + β colonies, with localized martensitic α′ observed at higher cooling rates. Estimated thermal gradients and solidification velocities enabled interpretation through a solidification-mode transformation framework, linking thermal history to grain morphology. The as-deposited clad hardness ranged from 385–455 VHN, gradually decreasing toward mid-height due to tempering from successive reheating, while heat-affected zone exhibited 355–380 VHN and the substrate retained ~340 VHN. Such gradients correlated with microstructural refinement and localized cooling conditions. Anisotropic tensile behavior, attributed to directional solidification and interlayer reheating, confirmed a microstructure–property relationship consistent with LMD-induced thermal cycling. Overall, the study highlights that scanning rate exerts dominant control over melt-pool geometry, grain refinement, and mechanical response, while inappropriate energy density promotes excessive coarsening.

Keywords:
Laser metal deposition; Ti-6Al-4V alloy; Laser output intensity; Feedstock mass flow; Melt pool dynamics.

1. INTRODUCTION

Laser–assisted additive fabrication has progressed into an essential manufacturing pathway intended for superior–integrity metallic assemblies and synonymous process labels including LMD (laser metal deposition), LSF (laser solid forming), DMD (direct metal deposition), and LENS (laser engineered net shaping) are regularly referenced to denote fabrication techniques, operating under principles of energy-assisted deposition [¹, ²]. Out of the listed technologies, LMD is receiving substantial industry focus owing to its ability to produce near–100% dense, near-design dimensionally accurate three-dimensional alloy architectures, built straight from design data. LMD depends on aligning a concentrated laser heat source simultaneously, with cladding material fed along beam axis, during which alloy powder is fed into the molten region, under controlled non-reactive gas atmosphere, to suppress unwanted oxidation and gas absorption [³]. The interaction between the laser and substrate creates a concentrated molten zone upon the base material or prior deposition, and as the laser beam progresses along the path, as instructed by pre-defined scanning coordinates, the melt pool undergoes rapid solidification, creating robust layer-to-layer bonding across the additive sequence of layers. Gradual layer integration under controlled temperature evolution facilitate the fabrication of structures possessing non-traditional geometrical profiles, which otherwise necessitate complex tooling and machining. Owing to advances in process optimization, robotic motion control, and alloy design, LMD has progressed into a competitive fabrication technique to support net-shape manufacturing, part repair, and functional modification of several metal alloys including alloys of Ti (like Ti-6Al-4V), applicable in space, medical prosthetics, and advanced mechanical design industries [⁴, ⁵].

The capability of LMD to fabricate near-net-shape titanium alloy components in a single processing step presents a transformative advantage for industries where low-volume, high-value metallic parts are required. By minimizing the dependence on conventional subtractive machining, tooling, and intermediate thermal processing, LMD offers substantial reductions in production lead time and overall manufacturing cost, particularly for alloys such as Ti-6Al-4V that are notoriously difficult and expensive to machine [⁶, ⁷]. In addition to its applicability in primary manufacturing, this technique is increasingly recognized for its effectiveness in component restoration and remanufacturing scenarios, where localized damage or wear can be rebuilt rather than replacing the entire component—an important advantage for aerospace and biomedical hardware where part availability and certification costs are high. Moreover, advancements in multi-material delivery systems, such as programmable multi-channel powder feeders and hybrid wire-powder deposition heads, enable the fabrication of graded structures with tailored composition and properties. These capabilities make LMD not only a viable additive manufacturing method for Ti alloys, but also a platform for producing engineered microstructures, performance-optimized functional surfaces, and next-generation gradient or hybrid material systems [⁸].

Despite the rapid progress of conventional Rapid Prototyping (RP) technologies, most polymer-based and low-melting-point metal processes still struggle to deliver components with the accuracy, mechanical strength, and surface integrity demanded for functional engineering applications. RP parts often remain limited to conceptual models or preliminary prototypes, requiring extensive auxiliary processing or remanufacturing steps to achieve service-ready performance [⁹, ¹⁰]. To overcome these constraints, the integration of RP principles with laser-based metal deposition has led to advanced directed energy deposition techniques capable of producing fully dense, structurally reliable metallic components directly from digital designs [¹¹]. In this context, Ti-6Al-4V emerges as a particularly important candidate for laser additive manufacturing due to its exceptional strength-to-weight ratio, corrosion resistance, and biocompatibility. However, its high reactivity, low thermal conductivity, and sensitivity to thermal gradients necessitate dedicated research to optimize LMD processing strategies. Accordingly, investigating the LMD behavior of Ti-6Al-4V is essential for enabling the direct fabrication of complex, high-performance titanium components for critical applications [¹²].

Although several research works on the LMD processing of Ti-6Al-4V alloys have provided valuable contributions, most of them focus on isolated aspects of the process, such as the influence of powder particle size and morphology, the formation of defects and surface characteristics, or the control of dimensional accuracy and layer quality [¹³,¹⁴,¹⁵,¹⁶]. Other studies have examined issues such as micro-hardness variations, tensile behavior, or localized corrosion performance of additively manufactured titanium components. However, these investigations generally do not present a unified or systematic experimental approach capable of comprehensively evaluating the microstructure, compositional uniformity, phase evolution, and mechanical response of as-deposited Ti-6Al-4V parts under varied LMD process conditions. As a result, the fundamental relationships between processing parameters, melt pool behavior, microstructural morphology, and resulting mechanical properties are still not fully established for titanium alloys [¹⁷, ¹⁸]. Furthermore, comprehensive experimental studies establishing robust processing–solidification–microstructure–property correlations for Ti-6Al-4V fabricated by LMD are still scarce.

To address these research gaps, the present investigation adopts a systematic orthogonal experimental design to evaluate the combined influence of laser output intensity, scanning rate, and feedstock mass flow over a wide processing window. Unlike previous studies that predominantly investigate individual process responses, this work integrates clad geometry optimization, melt-pool solidification analysis, microstructural characterization, hardness mapping, tensile anisotropy assessment, and thermal interpretation within a unified experimental framework. Furthermore, this work establishes quantitative correlations between deposition variables, estimated thermal conditions, solidification behavior, and the resulting mechanical response, enabling identification of an optimized deposition window for defect-free Ti-6Al-4V thin-wall structures. The findings provide a comprehensive process–microstructure–property relationship that contributes to improved scientific understanding and offers practical guidelines for process optimization of aerospace-grade Ti-6Al-4V components fabricated by laser metal deposition.

2. EXPERIMENTAL SETUP AND METHODOLOGY

2.1. Equipment configuration

The actual view of the Meltio M450 laser metal deposition (LMD) workstation used in this work is shown in Figure 1. The experimental platform integrates four primary subsystems: the laser energy source, multi-axis motion control unit, dual-feed material delivery system, and the digital control interface. The Meltio M450 employs a high-precision 1 kW multi-laser deposition head that combines six diode lasers to generate a concentrated energy beam suitable for processing titanium alloys. The integrated cooling system ensures thermal stability during prolonged deposition.

Figure 1
Actual view of Meltio M450 wire-and-powder Laser Metal Deposition (LMD) workstation used in this work.

The machine is equipped with a CNC-based three-axis motion control module, enabling precise manipulation of the build platform through coordinated x, y, and z-axis movements to construct three-dimensional components directly from sliced CAD data. The material delivery configuration includes both wire-feed and powder-feed capabilities, allowing flexible selection of feedstock depending on deposition requirements. A coaxial deposition nozzle ensures uniform feeding and stable melt pool formation, while the closed-loop control software manages laser power, deposition rate, and toolpath execution. This setup provides a robust and reliable manufacturing environment for fabricating Ti-6Al-4V samples under controlled process conditions.

The Meltio M450 wire-and-powder LMD workstation is operated through an integrated computer control system consisting of both hardware and software components. The control software converts CAD models into layer-wise toolpath instructions and manages all deposition operations through an intuitive graphical interface, enabling precise control of laser power, wire/powder feed, and motion sequencing. The hardware subsystem includes an industrial controller, servo motors, and a multi-axis motion control unit that coordinates the synchronized movement of the deposition head and build platform. These components function jointly to ensure accurate layer stacking and stable melt pool formation. The deposition process begins by establishing a stable thermal field on the Ti-6Al-4V substrate, followed by activation of the wire/powder feed to initiate continuous layer-by-layer fabrication. All subsystems operate cooperatively to achieve consistent melt pool stability and successful formation of three-dimensional components.

2.2. Material attributes and parameters

Ti-6Al-4V was used as the experimental material for the LMD process, employing gas-atomized alloy powder with a particle size range of 45–110 μm. Ti-6Al-4V is a widely used aerospace-grade alloy known for its high strength-to-weight ratio and excellent corrosion resistance, making it suitable for additive manufacturing. The substrate was also a Ti-6Al-4V plate having a dimension of 200 mm × 150 mm × 4 mm, and the chemical compositions of both powder and substrate (i.e., Ti) are provided in Table 1.

Table 1
Chemical constituents of research metal and substrate.

MATERIAL	Mo	Si	Nb	C	Sn	Al	Fe	V	Zr	Cr	Ti
Ti-6Al-4V alloy (Metal Powder & Substrate)	0.005	0.02	0.04	0.37	0.063	5.47	0.11	4.21	0.03	0.001	Balance

Based on the material characteristics of Ti-6Al-4V, research requirements, and equipment capabilities, the various LMD processing parameters are presented in Table 2.

Table 2
Input variables for LMD process.

PROCESS VARIABLES	OPERATING RANGE
Laser energy input	500–1500 W
Beam focal diameter	2 mm
Scanning velocity	3–15 mm/sec
Vertical increment amid layers	0.2–10 mm
Protective and transport gas	Argon
Delivery rate of protective gas	5–20 l/min
Feedstock mass flow	6–16 g/min
Nozzle to substrate spacing	12–20 mm

2.3. Deposition methodology

During the Laser Metal Deposition of Ti-6Al-4V, the selected processing parameters play a critical role in governing melt pool geometry, dilution characteristics, and interlayer bonding, which collectively determine the dimensional accuracy, microstructural evolution, and mechanical performance of the additively manufactured components. Therefore, identifying an optimal combination of LMD parameters with minimal experimental cost and time is essential. Given the complex and highly coupled nature of the LMD process, a systematic experimental strategy was adopted to evaluate the influence of key processing variables on deposition quality. An orthogonal experimental design was employed to determine the most significant parameters affecting the geometric characteristics and surface integrity of Ti-6Al-4V cladding layers. Primary process variables, including laser output intensity, scanning rate, and feedstock mass flow, were varied at multiple levels, and their effects on bead morphology and surface smoothness were analyzed. Furthermore, the orthogonal approach facilitated the quantitative assessment of parameter significance and interaction trends. Based on the experimental results, an optimized overlapping strategy was subsequently derived to achieve stable track formation and uniform layer deposition.

Using the optimized set of LMD parameters identified through the orthogonal experimental design, Ti-6Al-4V components were fabricated for detailed microstructural and mechanical characterization. The as-deposited samples were sectioned and prepared following standard metallographic procedures, and their microstructural features were examined using scanning electron microscopy (SEM). In addition, the thermal characteristics of the LMD process, including cooling rate and solidification velocity, were estimated through a combination of experimental measurements and analytical calculations. To elucidate the influence of processing parameters on microstructure evolution and mechanical behavior, deposits produced under different laser power and scanning speed conditions were systematically analyzed. Tensile specimens were extracted from the fabricated thin-wall structures, and mechanical testing was conducted to evaluate the strength and deformation characteristics of the LMD-fabricated Ti-6Al-4V alloy.

3. EXPERIMENTAL OUTCOMES AND INTERPRETATION

3.1. Verification of optimal deposition conditions

To successfully fabricate Ti-6Al-4V components using the LMD process, key processing parameters were systematically investigated and the resulting clad geometries were evaluated. Variations in parameters such as laser output intensity, scanning rate, and feedstock mass flow were found to significantly influence bead morphology, surface quality, and deposition stability. Therefore, experimental identification of an optimal parameter combination is essential to achieve defect-free, well-bonded deposits and to ensure the fabrication of high-quality Ti-6Al-4V components.

Laser Metal Deposition of Ti-6Al-4V involves complex interactions among multiple processing parameters, where even minor variations can significantly influence deposition stability, geometric accuracy, microstructural evolution, and mechanical performance of the fabricated components [⁴, ¹⁹]. To efficiently identify an optimal parameter combination while minimizing experimental effort, an orthogonal experimental design was employed based on commonly adopted LMD parameter ranges for titanium alloys. Laser output intensity, scanning rate, and feedstock mass flow were selected as the primary control factors, and each factor was assigned five representative levels. Accordingly, an L25 orthogonal array was constructed to systematically evaluate the effects of these parameters and their relative significance. This design enabled comprehensive assessment of parameter combinations through 25 deposition trials, providing statistically meaningful insights into the influence of individual factors on clad morphology and quality. The factors and their corresponding levels used in the orthogonal experiment are summarized in Table 3.

Table 3
Input variables and their assigned levels.

FACTOR	LASER OUTPUT INTENSITY, W	SCANNING RATE, mm/sec	FEEDSTOCK MASS FLOW, g/min
1st level	800	4	6
2nd level	950	6	8
3rd level	1100	8	10
4th level	1250	10	12
5th level	1400	12	14

For each of the 25 parameter sets, the resulting clad geometry was carefully measured, and key geometric characteristics such as clad height and clad width were recorded. Based on the results obtained from the orthogonal experiments, factor index graphs and variance analysis tables were constructed to evaluate the influence of individual processing parameters on clad geometry. Figure 2(a)–(c) and Figure 3(a)–(c) respectively illustrate the factor effect curves for clad height and clad width, where the horizontal axis represents the average values of clad geometry corresponding to each factor at different levels. These curves provide a clear visualization of how laser output intensity, scanning rate, and feedstock mass flow affect the deposition behavior of Ti-6Al-4V. Furthermore, the quantitative significance of each processing parameter was determined through analysis of variance, and the corresponding results are summarized in Tables 4 and 5.

Figure 2
Graphical illustration of impact of (a) laser output intensity (b) scanning velocity and (c) feedstock mass flow on height of the Ti-4Al-6V clads.

Figure 3
Graphical illustration of impact of (a) laser output intensity (b) scanning velocity and (c) feedstock mass flow on width of the Ti-4Al-6V clads.

Table 4
Variance analysis of clad width.

FACTOR	SUM OF DEVIATION SQUARE	DEGREE OF FREEDOM	F RATIO	F CRITICAL VALUE	SIGNIFICANCE
Laser output intensity	0.028	4	2.63	3.26	Not significant
Scanning rate	0.054	4	5.08	3.26	Significant
Feedstock mass flow	0.036	4	3.38	3.26	Very Significant
Error	0.032	12	–	–	–

Table 5
Variance analysis of clad height.

FACTOR	SUM OF DEVIATION SQUARE	DEGREE OF FREEDOM	F RATIO	F CRITICAL VALUE	SIGNIFICANCE
Laser output intensity	0.021	4	1.84	3.26	Not significant
Scanning rate	0.067	4	5.88	3.26	Very Significant
Feedstock mass flow	0.039	4	3.42	3.26	Significant
Error	0.034	12	–	–	–

As observed from factor index graphs and variance analysis tables, laser output intensity exhibits a relatively moderate influence on both clad height and clad width within the investigated parameter range. In contrast, scanning rate shows the most pronounced effect on the experimental results, indicating that it is the dominant factor controlling energy input per unit length and melt pool dynamics during Ti-6Al-4V deposition [²⁰].

The feedstock mass flow demonstrates an intermediate level of influence, affecting material supply efficiency and bead formation stability. Additionally, comparison of the results reveals that the effects of processing parameters on clad height are more significant than those on clad width, indicating that vertical material accumulation is more sensitive to parameter variations than lateral spreading of the melt pool.

Moreover, it can be observed from Figures 2 and 3 that when the laser output intensity exceeds a certain threshold, the clad height and width begin to decrease. This phenomenon can be attributed to excessive energy input, which leads to increased powder vaporization, spattering, and reduced powder utilization efficiency. For Ti-6Al-4V, excessive laser output intensity may also intensify thermal gradients and promote unfavorable microstructural evolution [²¹, ²²]. Therefore, while sufficient laser output intensity is essential to ensure stable melt pool formation and strong metallurgical bonding, excessively high laser power is not desirable, as it can adversely affect deposition quality, microstructure, and mechanical properties of the fabricated Ti-6Al-4V components.

Based on a comprehensive evaluation of the cladding results obtained from different orthogonal parameter combinations, an optimal set of LMD processing parameters for Ti-6Al-4V deposition was identified. The most favorable cladding quality was achieved at a laser output intensity of 1100 W, a scanning rate of 10 mm/s, and a feedstock mass flow of 8 g/min. Under these conditions, the deposited track exhibited stable melt pool behavior, continuous and uniform bead morphology, and a smooth surface finish without observable defects such as discontinuities or excessive spattering. Corresponding to this optimized parameter combination, the measured clad height and clad width were approximately 2.05 mm and 2.3 mm, respectively. These results indicate a balanced interaction between thermal input and material supply, ensuring efficient powder utilization and sound metallurgical bonding, thereby making this parameter set suitable for subsequent multilayer deposition and microstructural characterization of LMD-fabricated Ti-6Al-4V components [²³].

3.2. Optimal track overlap model

Single-track deposition experiments were conducted to determine the optimal scanning spacing, expressed in terms of overlapping percentage, for Ti-6Al-4V deposition by LMD. The degree of overlap plays a crucial role in controlling surface flatness and dimensional accuracy of the deposited layers. An inappropriate overlapping percentage may lead to surface waviness, local height variation, or accumulation of thermal distortion. Under such conditions, the effective energy input and material deposition per pass become non-uniform, thereby altering melt pool stability and solidification behavior. Consequently, improper overlap selection can adversely affect geometric consistency and may even result in deposition failure during multilayer fabrication of Ti-6Al-4V components [²⁴].

The metallographic cross-section of a single deposited track, as illustrated in Figure 4(a), reveals that the cladding profile of Ti-6Al-4V can be reasonably approximated as a segment of a circular geometry. In this representation, H denotes the height of an individual cladding layer, while W represents the corresponding clad width. Based on this geometric approximation, the ideal overlapping condition between two adjacent deposition tracks is schematically shown in Figure 4(b). Under the ideal overlap state, adjacent tracks exhibit uniform height and a smooth interfacial transition, thereby eliminating inter-track porosity and ensuring continuous metallurgical bonding between successive layers. This idealized geometric model provides the basis for calculating the optimal overlapping percentage required for stable multilayer deposition of Ti-6Al-4V components [²⁵].

Figure 4
(a) Metalograph of the cross section of single pass cladding and (b) schematic of the cross section in the ideal overlapping state.

In the proposed ideal overlapping model, two neighboring Ti-6Al-4V cladding tracks are assumed to possess identical cross-sectional geometry, such that both the maximum cladding height and the cross-sectional area remain equal. Here, W denotes the width of a single deposited track, while H represents the height of two adjacent clads. The cross-section of the former track can be approximated as a circular arc C₁, whereas the overlapping portion of the subsequent track can be described by a corresponding circular arc C₂, both having an identical radius R. Under ideal conditions, the surface formed between C₁ and C₂ is planar, which is essential for suppressing pore formation and ensuring uniform metallurgical bonding between neighboring tracks and successive layers. In this geometric configuration, the concave groove area (A₁) formed between adjacent tracks is exactly compensated by the overlapping area (A₂) of the subsequent deposit, such that A₁ = A₂. This balance allows the overlapping material to completely fill the inter-track depression, resulting in a smooth and continuous clad surface. Accordingly, the overlapping percentage ηc and the scanning spacing (S) corresponding to this condition are defined as optimal for stable multilayer deposition of Ti-6Al-4V components and must be quantitatively determined. Based on the geometric relationship illustrated in Figure 5, the center of the circular arc C₁ corresponding to a Ti-6Al-4V single-track cross section is located at (0, H−R), and its governing equation can be described as

Figure 5
Ti-6Al-4V alloy clads fabricated using LMD process.

(1) $$y=(H-R)+\sqrt{R^2+x^2}$$

(2) $$\text{where}R=\frac{W^2}{8H}+\frac{H}{2}$$

When y ≥ 0, Equation (1) can be modified to the below mentioned form:

(3) $$y-H=\sqrt{R^2-x^2}-R$$

Accordingly, the concave groove area (A₁) and the overlapping area (A₂) associated with adjacent Ti-6Al-4V cladding tracks were evaluated using an integral-based geometric approach [²⁶]. The corresponding calculation expressions are formulated as follows:

(4) $$A_1=S\left(R-\frac{1}{\text{2}}\sqrt{R^2-{\left(\frac{s}{2}\right)}^2}\right)-R^2\text{arcsin}\left(\frac{s}{2R}\right)$$

(5) $$A_2=R^2\left[\text{arcsin}\left(\frac{W}{2R}\right)-\text{arcsin}\left(\frac{s}{2R}\right)\right]+\frac{1}{\text{2}}\left[W\sqrt{R^2-{\left(\frac{W}{2}\right)}^2}-S\sqrt{R^2-{\left(\frac{S}{2}\right)}^2}\right]+\left(H-R\right)\left(W-S\right)$$

From the geometric relationship illustrated in Figure 4(b), the scanning spacing S varies continuously between 0 and the clad width W. When S = 0, complete overlap occurs and the overlapping percentage is ηc = 100%. Conversely, when S = W, there is no overlap and ηc = 0. On this basis, the overlapping percentage for Ti-6Al-4V cladding can be expressed by the following general relationship:

(6) $$S=W\left(1-\frac{\eta c}{100}\right)$$

In practice, the determination of the optimal overlapping percentage for Ti-6Al-4V deposition is achieved by iteratively adjusting the scanning spacing to obtain the optimal value S, under the condition that A₁ = A₂ [²⁷]. For this purpose, the clad width W is discretized into a sufficiently large number of segments n, and a small tolerance ε is prescribed. The scanning spacing is then defined as S = (W/n)* i, where i ∈ [0, n], and the optimal overlap is identified when the condition |A₁ − A₂| < ε is satisfied [²⁸].

Accordingly, once the clad height H and clad width W are experimentally measured for Ti-6Al-4V single-track deposits, the optimal scanning spacing S under ideal overlapping conditions can be determined using the proposed optimization algorithm. The corresponding optimal overlapping percentage ηc is subsequently calculated using Equation (6). It should be noted that the above derivations are based on an idealized geometric assumption [²⁹]. In practice, due to the influence of surface tension and melt pool dynamics inherent to Ti-6Al-4V, the overlapping surface is rarely perfectly planar, and slight convexity at the track crest or concavity between adjacent tracks may occur. This effect becomes more pronounced for deposits with larger clad heights, where increased curvature can adversely affect surface uniformity. Therefore, excessive clad height should be avoided to ensure stable deposition and improved forming quality [⁷, ¹⁸].

3.3. Single-layer deposition investigation

Based on the optimal clad geometry obtained under the favorable deposition conditions identified through the orthogonal experimental design, the measured average clad height and width for Ti-6Al-4V were H ≈ 2.05 mm and W ≈ 2.3 mm, respectively. These values were incorporated into the ideal overlapping model to determine the optimal scanning spacing and corresponding overlapping percentage, yielding S ≈ 1.56 mm and ηc ≈ 32%. To validate the feasibility of the proposed experimental methodology, planar scanning experiments were conducted using the optimized processing parameters (laser output intensity of 1100 W, scanning rate of 10 mm/s, and feedstock mass flow of 8 g/min) in conjunction with the calculated scanning spacing. Photograph of the laser metal deposited Ti-6Al-4V clads is illustrated in the Figure 5. The deposited surface exhibits uniform track spacing, smooth morphology, and absence of visible cracks or porosity, confirming the effectiveness and reliability of orthogonal design and ideal overlapping model for Ti-6Al-4V clad deposition.

4. CHARACTERIZATION OF BONDING, MICROSTRUCTURE, AND HARDNESS

4.1. Interfacial bonding layer

Figure 6 illustrates the microstructural features at the interface between the deposited Ti-6Al-4V clad and the substrate. The cross-section clearly shows the transition from the substrate to the interfacial bonding layer and then to the deposited clad. A thin, continuous interfacial layer is evident, formed by localized melting and alloy inter-diffusion under the thermal field of the deposition process. This layer appears slightly brighter after etching, and exhibits a refined, planar-to-columnar transition structure, indicating epitaxial growth from the substrate toward the clad [³⁰]. Immediately above the interface, elongated prior-β grains and basket-weave (α + β) colonies develop, while the upper region of the clad shows progressively coarser, directional structures ­consistent with solidification under steep thermal gradients. The continuity of the interfacial layer and the absence of cracks or un-melted regions confirm strong metallurgical bonding, providing an effective load-transfer path between the substrate and the deposited Ti-6Al-4V clad [³¹].

Figure 6
Microstructural image highlighting the metallurgical bond and directional growth in Ti-6Al-4V clad.

In accordance with solidification theory, the ratio between the local temperature gradient (here denoted as T) and the solidification velocity (denoted as SV) governs the selection of interfacial morphologies during LMD. A high T/SV condition favors a stable planar front, whereas moderate reductions promote columnar growth; when T/SV becomes small, equiaxed structures dominate [²¹, ³²]. In the present Ti-6Al-4V deposits, as seen in the Figure 6, the region immediately adjacent to the interfacial bonding layer shows a thin planar zone formed under steep thermal gradients at the melt boundary. Moving upward, T decreases while SV increases, leading to the evolution of oriented columnar prior-β grains, followed by basket-weave (α + β) colonies within the clad. This graded transition confirms that the interfacial layer solidified under highly directional heat flow, while the upper clad experienced progressively less constrained solidification [¹¹, ²²].

In the present study, the thermal field established during deposition produces a pronounced gradient near the base of the melt pool. At the immediate vicinity of the substrate, T/SV is relatively large, stabilizing the solid–liquid interface and promoting the formation of a thin, feature-poor interfacial layer that transitions gradually into the clad. With increasing distance from the substrate, T/SV decreases rapidly as heat dissipates into the surrounding material, and the planar interface becomes unstable. Consequently, competitive grain growth governs the solidification, leading to elongated columnar features that propagate upward from the interfacial region toward the clad interior [³³]. The morphology observed in Figure 6 is therefore characteristic of epitaxial solidification, where the advancing solid phase continuously inherits crystallographic orientation from the underlying substrate. Because the vertical component of T dominates within most of the melt pool, grains oriented closest to this preferential direction grow more rapidly, progressively out-competing mis-oriented neighbors. This directional selection yields the columnar architecture visible across the clad thickness, providing metallurgical continuity across the interfacial bonding layer and ensuring effective load transfer between the substrate and the deposited Ti-6Al-4V clad.

In addition, the preferred grain growth direction in the vicinity of the interfacial bonding layer is slightly inclined relative to the local normal of the solid–liquid boundary, reflecting the directional solidification characteristics of LMD under steep temperature gradients (T) and comparatively high solid–liquid interface velocities (SV). Frequent heterogeneous nucleation ahead of the advancing solidification front increases the nucleation density, thereby promoting significant microstructural refinement. Guided by the thermal gradient and crystallographic orientation inheritance, the clad developed epitaxial prior-β columnar grains extending from the substrate–clad interface, followed by transformation into refined α + β colonies during cooling. This microstructural evolution is in good agreement with earlier studies on LMD-fabricated Ti-6Al-4V alloys, which similarly reported epitaxial β-columnar growth driven by directional heat extraction and subsequent α + β phase transformation during rapid solidification [¹⁸, ³⁴]. However, unlike these studies, which primarily described the evolution of such microstructures qualitatively, the present investigation systematically correlates the observed β-columnar growth, α + β refinement, and microstructural homogeneity with variations in laser output intensity, scanning rate, and feedstock mass flow through their influence on thermal history and solidification behavior. Consequently, the present work not only validates previously reported solidification mechanisms but also establishes a comprehensive processing–solidification–microstructure–property relationship for LMD-fabricated Ti-6Al-4V clads, thereby providing a more complete understanding of deposition-variable-dependent microstructural evolution and its influence on mechanical performance.

4.2. Macroscopic structure grain-morphology prediction diagram

For the Ti-6Al-4V clad produced under optimized deposition conditions, the local solidification environment within the melt pool had promoted dendritic growth as the dominant primary morphology, while a thin planar layer formed adjacent to the substrate. From the bottom toward the top of the melt pool, T decreases gradually, whereas SV increases progressively, resulting in a continuous distribution of solidification conditions. These evolving parameters were superimposed on the solidification-mode transformation (SMT) Curve (i.e., a curve illustrating the transformation from columnar to equiaxed morphology) to construct the grain-morphology prediction diagram shown in Figure 7. The shaded region represents the feasible T–SV window encountered during LMD of Ti-6Al-4V under optimum conditions [³⁵]. Most of this domain lies on the columnar side of the SMT curve, explaining the predominance of epitaxial columnar grains, while localized excursions toward lower T/SV conditions promote partial formation of equiaxed grains. This map therefore provides a quantitative framework linking melt-pool thermal fields to macroscopic grain morphology in Ti-6Al-4V. The methodology used to determine the Ti-6Al-4V T–SV envelope is elaborated in Section 5.1.

Figure 7
Grain-morphology prediction diagram illustrating diffusion bonding and columnar growth in Ti-6Al-4V clad.

As shown in Figure 7, the solidification path corresponding to the optimized deposition parameters for Ti-6Al-4V predominantly lies within the columnar-growth domain, indicating that SMT does not occur under these conditions. This prediction agrees well with the metallographic observations in Figure 6, where the clad largely exhibits epitaxial columnar grains extending from the substrate toward the top surface, with only limited evidence of equiaxed refinement. The close consistency between model prediction and experiment confirms the reliability of the transition framework adopted in this study [¹⁸, ²⁶].

In contrast to certain nickel-based alloys reported in the literature [⁴, ¹⁸], the combination of relatively high T and moderate SV in Ti-6Al-4V maintains the melt-pool trajectory to the right of the SMT boundary, preventing full development of equiaxed grains at the clad crown. Nevertheless, the gradual decrease in T and concurrent rise in SV from the melt-pool base upward reveals that careful adjustment of process parameters can deliberately shift the thermal path across different microstructure regions. Therefore, the Ti-6Al-4V grain-morphology prediction diagram demonstrates that tailored control of T and SV provides an effective tool to engineer grain morphology—ranging from strongly epitaxial columnar growth to partially or fully equiaxed structures—depending on application requirements.

4.3. Microstructural morphology

Figure 8(a) and (b) illustrates the representative microstructural features across distinctive interior regions of the Ti-6Al-4V clad. The cross-section reveals a predominantly dendritic/columnar morphology that grows epitaxially from the substrate toward the clad surface, consistent with directional solidification under a steep temperature gradient. The overall geometry of successive deposited tracks exhibits a pronounced Non Planar Interface, originating from the combined effects of melt pool shape, wetting behavior, and surface-tension-driven solidification dynamics [²³].

Figure 8
Cross-sectional micro-structural images showing internal morphology of Ti-6Al-4V clad at (a) left-side and (b) right-side positions.

In the vicinity of the interface, relatively coarse primary dendrites (corresponding to prior-β grains) are observed, accompanied by well-developed secondary arms. Moving upward through the clad, the dendrite arm spacing progressively decreases and the structure becomes noticeably refined, reflecting the enhanced cooling rate and reduced local solidification time near the upper region of the melt pool. For Ti-6Al-4V, this refinement promotes the formation of fine α/α′ laths within the transformed prior-β grains, producing a dense and homogeneous microstructural network.

The observed gradient in dendrite/coarseness can be attributed to the thermal history of successive layers: material near the base experiences reheating and partial tempering due to multiple thermal cycles, while newly solidified regions near the top cool rapidly under the action of shielding gas and environmental heat extraction. Throughout the clad, microstructural defects such as porosity and inclusions are scarce. This can be ascribed to vigorous melt-pool convection induced by the laser, which enhances fluid flow and solute redistribution, enabling entrapped gases and inclusions to rise and escape, thereby ensuring good metallurgical integrity of the Ti-6Al-4V clad [²⁰, ³⁶].

4.4. Elemental distribution and phase constitution

Figure 9 illustrates the equilibrium phase relations of the Ti-6Al-4V alloy system. The composition path from the clad region toward the substrate reflects gradual dilution of Al and V, thereby modifying the local β-stabilizing and α-stabilizing tendencies. Along this trajectory, the phase constitution evolves from a predominantly α + β field within the clad to regions progressively enriched in α, consistent with the equilibrium boundaries indicated in the diagram. This progressive shift implies that the local solid-state transformations during cooling are governed mainly by the reduction in β-stabilizer content, which narrows the stability domain of the β phase and promotes α precipitation, ultimately defining the phase constitution across the deposited layer [⁶, ¹¹].

Figure 9
Equilibrium phase relations of the Ti-6Al-4V alloy system.

To quantify elemental redistribution in the clad, compositional measurements were performed along directions crossing adjacent prior-β dendrite arms at regularly spaced locations. The selected points capture both dendrite cores and inter-dendritic regions, enabling assessment of spatial variation at the melt solidification scale. The corresponding concentration profiles of Ti, Al and V obtained from these points are plotted in Figure 10, while Figure 11 presents a representative EDS (Energy-dispersive X-ray spectrum) spectrum from the 1st measurement position. As observed in Figure 10, the variations in elemental content are relatively small, and the average composition closely matches the nominal Ti-6Al-4V powder. Mild enrichment of V at inter-dendritic regions and slightly higher Al content near dendrite cores are evident, consistent with the β-stabilizing tendency of V and the α-stabilizing effect of Al. The calculated segregation ratios for Ti, Al and V indicate only limited micro-segregation, confirming that the deposited clad exhibits near-homogeneous composition at both macro- and micro-scales [³⁷].

Figure 10
Elemental concentration fluctuations of Ti, Al and V within Ti-6Al-4V melt-solidified microstructure.

Figure 11
Elemental identification by EDS (Energy-dispersive X-ray spectrum) at 1st analyzed point within the Ti-6Al-4V clad layer.

This observation is rationalized by the rapid solidification and high thermal gradient intrinsic to LMD, which restrict long-range solute partitioning and promote subsequent partial homogenization through layer-overlap reheating [⁵, ¹⁴]. Consequently, elemental diffusion between dendrite cores and inter-dendritic regions is enhanced, suppressing solute pile-up and reducing segregation severity compared with conventionally cast Ti-6Al-4V. The resulting compositional uniformity supports formation of a refined α/β microstructure and minimizes the likelihood of composition-induced phase localization, contributing favorably to the mechanical stability of the clad [²³].

Figure 12 presents the X-ray diffraction profile of the Ti-6Al-4V clad, clearly revealing the coexistence of dominant α-Ti (hcp) reflections together with a weak β-Ti (bcc) peak. The principal diffraction maxima corresponding to the α phase — indexed as α(100), α(101), and α(102) — indicate that solidification proceeded primarily through transformation to the equilibrium α matrix during cooling.

Figure 12
X-ray diffraction profile of the fabricated Ti-6Al-4V clad.

A minor β (110) peak persists at higher 2θ, implying retention of a small fraction of metastable β phase, which is typical in rapidly solidified LMD builds due to incomplete diffusion-controlled partitioning of Al (α-stabilizer) and V (β-stabilizer). Compared with conventional wrought Ti-6Al-4V, the relative intensities of α peaks suggest mild crystallographic texture evolution associated with directional heat flow. The absence of secondary intermetallic or hydride signatures confirms that compositional dilution, oxidation, and hydrogen pickup were effectively controlled during processing. Overall, the diffraction response substantiates that the clad develops a predominantly α-Ti matrix with limited retained β, consistent with the solidification path predicted from equilibrium phase diagram and high cooling rates inherent to laser metal deposition [²⁵, ³⁸].

4.5. Hardness gradient across deposition interface

To assess the mechanical response of the deposited structure, the Vickers micro-hardness was measured across the clad–substrate cross-section. Table 6 lists the comparison between the as-deposited Ti-6Al-4V clad and the wrought substrate. As seen from the table, the clad exhibits consistently higher hardness, which is attributed to rapid solidification, grain refinement, and a fine α + β matrix formed during LMD. A smooth hardness transition is observed toward the interface, indicating effective metallurgical bonding and the absence of abrupt property discontinuities—an essential feature for resisting service-induced stress concentrations.

Table 6
Observed and standard hardness values w.r.t Ti-6Al-4V.

ALLOY	OBSERVED HARDNESS (VHN)	NOMINAL/STANDARD HARDNESS (VHN)
Mill annealed Ti-6Al-4V plate	–	340 VHN
Heat-affected region (near interface) of as-deposited Ti-6Al-4V clad	355–380	–
As-deposited Ti-6Al-4V clad	385–455	–

The hardness gradient across the deposition interface of the Ti-6Al-4V build is summarized in Figure 13 and three salient features can be observed from this Figure. First, the Vickers Hardness Number (VHN) at the clad surface is distinctly higher than within the clad interior, attributed to the strong convective cooling from the shielding gas that promotes rapid solidification and grain refinement. Second, the region near the interface exhibits slightly elevated hardness (≈ 355–380 VHN) compared with the bulk substrate, reflecting the combined effects of epitaxial solidification and partial solution strengthening. Third, a mild softening trend appears within the substrate adjacent to the melt pool, resulting from thermal cycling and localized tempering during LMD. Overall, the clad maintains the highest hardness (≈ 385–455 VHN), outperforming the mill-annealed substrate (~ 340 VHN) and evidencing the beneficial refinement associated with rapid solidification [²⁴].

Figure 13
Variation of Vickers hardness along the build height of as-deposited Ti-6Al-4V clad joint.

5. HEAT DISSIPATION, SOLIDIFICATION RATE AND STRUCTURAL INTEGRITY

The preceding results indicate that the macroscopic morphology and grain evolution in LMD-processed Ti-6Al-4V are governed primarily by the ratio between the local temperature gradient (T) and the solidification velocity (SV). Hence, a detailed assessment of heat dissipation and solidification rate was undertaken. In this investigation, clad tracks produced with identical processing conditions—laser power of 1100 W, powder feed rate of 8 g/min, and spot diameter of 2 mm—but different scanning speeds were examined. Two representative scanning speeds, 4 mm/s and 10 mm/s, were selected to quantify the variation in T, SV, and the resulting solidification behavior across the deposited layers.

5.1. Heat dissipation intensity

To quantify T and SV, the heat-dissipation intensity must first be established. In LMD solidification, the cooling history is reliably inferred from the secondary dendrite arm spacing (SDAS). For Ti-6Al-4V, well-validated empirical correlations between cooling rate (i.e., ε) and SDAS (i.e., λ₂) are available [³⁹] and were applied to deposits produced at 1100 W, 8 g/min, 2 mm spot size, and scanning speeds of 4 and 10 mm/s. For each clad, a representative cross-section was etched and the SDAS (λ₂) was measured along the melt-pool depth. For Ti-6Al-4V, the cooling rate (ε) is empirically related to λ₂ by a power law [³⁴, ³⁹] of the form:

(7) $$\epsilon ={\left(\frac{C}{{\lambda }_2}\right)}^{1/n}$$

where C and n are material-specific constants determined from prior titanium alloy solidification studies and validated against our measurements. The cooling rate ε is reported in kelvin per second, while λ₂ is represented in micrometers (μm), consistent with prior Ti-6Al-4V studies [¹³, ¹⁸]. In order to quantify the heat-dissipation intensity, representative cross-section scanning electron microscopic (SEM) images were examined for Ti-6Al-4V clads fabricated at scanning speeds of 4 mm/s and 10 mm/s, and are illustrated as Figure 14(a) and (b) respectively.

Figure 14
SEM images of Ti-6Al-4V clads fabricated at scanning speeds of (a) 4 mm/s and (b) 10 mm/s.

Similar to the reference case, the Ti-6Al-4V clads reveal a directional solidification morphology, where the prior-β growth direction is decorated by refined α platelets. The regions marked as red colored circles correspond to the secondary dendrite arms, from which the secondary dendrite arm spacing (λ₂) was extracted using the micro-scale in the figure.

For the lower scanning speed (Figure 14(a)), λ₂ is typically 5–6 μm, whereas at the higher scanning speed (Figure 14(b)) it reduces to 2–3 μm, indicating more efficient thermal extraction from the melt pool. Using the Equation (7), λ₂ = Cε − n (with C ≈ 45–55 and n ≈ 0.30–0.35), the corresponding heat-dissipation intensities for Ti-6Al-4V clads are obtained as ε∼103 K/s for 4 mm/s and ε∼104 K/s for 10 mm/s. These values are consistent with the rapid solidification regime typically reported for laser-processed Ti-6Al-4V. From a processing standpoint, the powder addition rate, scanning speed and laser output must remain mutually balanced to avoid instability of the melt pool [⁴⁰]. Consequently, excessive increases in scanning speed require proportionate adjustment of laser energy input to sustain full densification. Within the feasible parameter envelope for Ti-6Al-4V, the heat-dissipation intensity can therefore be expected to vary over roughly 10²–10⁴ K/s, directly governing the refinement of α morphology and, ultimately, the hardness gradient across the deposition interface.

5.2. Solid–liquid interface velocity

In laser metal deposition of Ti-6Al-4V, the solid–liquid interface velocity (SV) represents the true growth rate of the solidification front along the direction of heat extraction. Similar to laser surface re-melting analyses, SV can be inferred geometrically from the inclination angle (θ) between the columnar growth orientation and the beam-scanning direction through Equation (8):

(8) $$\theta =Co{s}^{-1}\left(\frac{S_V}{S_r}\right)$$

where S_r is the rate of scanning. From metallographic observations in the present work, the dominant growth orientation is inclined by θ ≈ 20–25° to the scanning direction, reflecting strong alignment with the heat-flow vector [¹², ³⁵]. Using the two process conditions investigated (4 and 10 mm/s), the resulting interface velocities are approximately SV ≈ 3.4–3.9 mm/s and SV ≈ 8.9–9.3 mm/s, respectively. These magnitudes are consistent with rapid directional solidification typically reported for Ti-6Al-4V LMD deposits. Because the cooling rate (ε) relates to thermal gradient (T) through ε = T⋅SV, combining these SV values with the experimentally inferred cooling-rate range yields thermal gradients on the order of 1×106 to 2×106 K/m. Although such estimates rely on geometric assumptions and empirical correlations, they convincingly indicate a regime of high interface velocity and steep thermal gradients—conditions that promote fine basket-weave α/β morphology and suppressed coarsening in the deposited Ti-6Al-4V clads [⁴¹].

5.3. Performance attributes of fabricated samples

To evaluate the performance attributes of the LMD-fabricated Ti-6Al-4V samples, tensile specimens were extracted from the deposited walls using precision wire-cutting. The schematics of the specimen geometry and orientation with respect to the laser scanning path are illustrated in Figure 15(a)–(c). Two tensile directions (parallel and perpendicular to the scan tracks) were considered to assess anisotropy. Mechanical testing revealed stable load–displacement behavior and consistent failure responses across both orientations, indicating good metallurgical bonding and structural integrity throughout the deposited regions.

Figure 15
Geometry and orientation of tensile specimens prepared from laser-deposited Ti-6Al-4V: (a) reference geometry, (b) loading parallel to build/scan direction, (c) loading perpendicular to build/scan direction.

The corresponding fracture morphologies are presented in Figure 16(a)–(b). Both specimens exhibited rough, dimple-dominated fracture surfaces with no evidence of cleavage, confirming a ductile failure mechanism. The uniformly distributed dimples indicate effective metallurgical bonding and adequate plastic accommodation across layers, demonstrating that the LMD-processed Ti-6Al-4V maintains sound toughness and structural integrity suitable for load-bearing applications [¹², ³¹]. As seen in Figure 16(a)–(b), the fracture surfaces exhibit a rough appearance with numerous equi-axed and elongated dimples, confirming a predominantly ductile failure mode and indicating that the LMD-fabricated Ti-6Al-4V samples retain good damage tolerance. The corresponding mechanical results are summarized in Table 7.

Figure 16
SEM fractography of LMD-deposited Ti-6Al-4V clad tensile specimen with loading axis (a) parallel and (b) perpendicular to the scanning direction.

Table 7
Comparison of mechanical properties of LMD Ti-6Al-4V thin-wall parts with conventional material.

ALLOY	TENSILE STRENGTH, MPa	YIELD STRENGTH, MPa	ELONGATION, %
As-deposited part Ti-6Al-4V Clad (Perpendicular)	1008	928	10.6
As-deposited part Ti-6Al-4V Clad (Parallel)	962	885	11.8
Mill-annealed Ti-6Al-4V plate	930	870	13.5

It can be observed from Table 7 that the yield and tensile strengths of the LMD-fabricated Ti-6Al-4V thin-wall components are slightly higher than those of the mill-annealed Ti-6Al-4V plate, whereas the elongation of the deposited material is comparatively lower. The enhanced strength of the LMD-fabricated parts can be attributed to the rapid solidification during the deposition process, which results in the formation of a fine acicular α′ martensitic microstructure [⁴²]. However, the ductility is reduced due to the presence of martensitic phases and residual stresses generated during the layer-by-layer deposition process.

To quantitatively assess the directional dependence of the mechanical response, the tensile properties obtained in the two loading orientations were compared as seen in Table 8. The specimen loaded parallel to the scanning direction exhibited a tensile strength of 1008 MPa, representing approximately 4.8% higher strength than that measured in the perpendicular direction (962 MPa). Likewise, the yield strength increased from 885 MPa to 928 MPa, corresponding to an enhancement of approximately 4.9%. Conversely, the elongation increased from 10.6% in the parallel orientation to 11.8% in the perpendicular orientation, indicating an improvement in ductility of approximately 11.3%. These quantitative differences confirm the presence of moderate mechanical anisotropy arising from the directional solidification characteristics, epitaxial grain growth, and cyclic thermal history inherent to the layer-by-layer laser metal deposition process.

Table 8
Quantitative assessment of anisotropic tensile behavior of LMD-fabricated Ti-6Al-4V.

PROPERTY	PARALLEL	PERPENDICULAR	DIFFERENCE (%)
Tensile Strength (MPa)	1008	962	4.8
Yield Strength (MPa)	928	885	4.9
Elongation (%)	10.6	11.8	11.3

In addition, the tensile response of the LMD-fabricated Ti-6Al-4V specimens exhibits clear anisotropy with respect to the scanning direction. As shown in Figure 16(a), specimens loaded parallel to the scanning direction develop comparatively higher strengths but fail preferentially along interlayer regions, where incomplete metallurgical bonding can act as crack initiation sites. In contrast, the specimens tested perpendicular to the scanning direction (Figure 16(b)) show slightly reduced strength but noticeably higher elongation, consistent with the dimpled, deeper fracture morphology indicating superior ductility along the primary growth orientation. This directional dependence arises from the inherently layered nature of laser deposition [¹⁵, ²³]. Depending on the intended service conditions, such anisotropy may be either exploited or mitigated through suitable post-deposition heat treatments, such as solutionizing followed by aging.

To elucidate the hardness response of the Ti-6Al-4V thin-wall samples, a systematic micro-hardness mapping strategy was adopted. Representative indentation locations, shown in Figure 17, were selected to quantify hardness variations along both the build height and wall width.

Figure 17
Layout of hardness test positions within the cross-section of the Ti-6Al-4V thin-wall produced by LMD, indicating re-melting zones and layer-wise spacing of indentation marks.

The spacing between successive test points was maintained at 0.2 mm in the vertical direction and 0.1 mm in the transverse direction, with the first indentation positioned 0.2 mm below the top surface and indexed as Point 1. Subsequent points (2–12) were placed downward at equal intervals, while ten additional indentations were symmetrically distributed about the central region in the width direction.

As illustrated in Figure 18(a), the hardness distribution across the width of the Ti-6Al-4V thin-wall clad exhibits a characteristic “edge-hardened” profile, where the values at both sides remain higher than those near the central region. This trend is attributed to intensified convective and conductive heat extraction from the exposed edges, which promotes faster cooling and the formation of finer martensitic α′ structures, thereby enhancing hardness [⁶, ⁴³]. In contrast, Figure 18(b) reveals a progressive reduction in hardness from the upper layers toward the mid-height of the wall, followed by a softened zone near the substrate. This behavior reflects the cumulative tempering imposed by successive layer deposition: reheating of previously solidified tracks coarsens the microstructure and partially decomposes α′ into α + β, reducing hardness [¹⁸]. Meanwhile, the top region experiences rapid cooling under the combined influence of ambient air and shielding gas, without significant re-melting, and therefore retains the highest hardness.

Figure 18
(a) and (b) Hardness profiles of the LMD-fabricated Ti-6Al-4V thin-wall structure measured in (a) width direction and (b) height direction.

The locally reduced hardness near the re-melting bands (identified adjacent to the layer interfaces) corresponds to regions subjected to slower cooling and extended thermal exposure. A similar trend was observed in monolayer Ti-6Al-4V clads (Figure 13), although the multilayer structure shows moderately lower hardness overall, confirming the dominant role of repeated thermal cycling in diminishing the cooling-induced hardening effect [²⁴, ⁴³]. In practical terms, the harder surface layers imply improved wear and corrosion resistance, whereas the comparatively softer interior imparts beneficial ductility and damage tolerance. Thus, the spatial hardness gradients inherent to LMD processing can be strategically leveraged or mitigated through optimization of laser parameters and tailored post-heat-treatments, depending on the functional performance required for aerospace-grade Ti-6Al-4V components.

6. CONCLUDING REMARKS

Based on the comprehensive LMD experiments with Ti-6Al-4V powder, the heat-dissipation intensity and solid–liquid interface velocity in the deposition process were evaluated, and the influencing regularities of critical deposition variables on the geometric profile, microstructure and mechanical response were systematically analyzed. As a result, the conclusions can be drawn as follows:

• The orthogonal design strategy combined with the ideal overlapping model enabled rational selection of deposition variables. Using the optimal combination — laser output intensity of 1100 W, scanning rate of 10 mm/s, and feedstock mass flow of 8 g/min — continuous thin-wall structures were produced without visible pores, cracks, or lack-of-fusion defects, demonstrating the reliability of the optimized LMD process for Ti-6Al-4V.

• Metallographic examination confirmed a strong solution-bonding layer formed through alloy inter-diffusion at the substrate–clad interface. The clad exhibited epitaxial grain growth from the substrate, followed by directional columnar-to-refined basket-weave α + β morphology, governed by combined effects of temperature gradient (G) and solid–liquid interface velocity (Sv).

• The cooling behavior revealed that higher scanning rate and reduced effective heat input increased the heat dissipation intensity, promoting finer prior-β grain size and reduced α-lamella thickness. Rapid solidification minimized segregation and suppressed secondary dendrite coarsening. Consequently, the refined microstructure enhanced hardness and strength, while maintaining metallurgical continuity across layers.

• Tensile results demonstrated mechanical strengths comparable to conventionally processed Ti-6Al-4V, accompanied by ductile fracture characterized by dense dimples. Hardness mapping showed elevated values near outer surfaces, moderate reduction in mid-height regions due to tempering effects, and locally softened zones near reheated interfaces — trends consistent with multilayer thermal cycling.

• The study establishes that proper balancing of laser output intensity, scanning rate, and feedstock delivery is critical to stabilize melt-pool dynamics, control solidification conditions, and minimize anisotropy. Optimized heat management simultaneously prevents grain coarsening, improves interlayer bonding, and yields mechanically consistent components.

Overall, the findings provide a practical framework for tailoring LMD parameters for aerospace-grade titanium structures, while offering a foundation for future work on predictive modeling, hybrid heat-treatment strategies, and multi-layer integrity assessment in high-performance Ti-6Al-4V builds.

7. DATA AVAILABILITY

All data that support the findings of this study are included within the article (and any supplementary files).

8. BIBLIOGRAPHY

• ^[1] BAGHDADCHI, A., CARY, C., SRIDHAR, N., et al, “Corrosion resistance and microstructure analysis of additively manufactured 22% chromium duplex stainless steel by laser metal deposition with wire”, Journal of Materials Research and Technology, v. 26, pp. 6741–6756, 2023. doi: https://doi.org/10.1016/j.jmrt.2023.09.037.
  » https://doi.org/10.1016/j.jmrt.2023.09.037
• ^[2] BANJANIN, B., VLADIC, G., PÁL, M., et al, “Consistency analysis of mechanical properties of elements produced by FDM additive manufacturing technology”, Matéria (Rio de Janeiro), v. 23, n. 4, pp. e12250, 2018. doi: https://doi.org/10.1590/s1517-707620180004.0584.
  » https://doi.org/10.1590/s1517-707620180004.0584
• ^[3] YAMASHITA, Y., KAMO, S., SATO, Y., et al, “Laser-assisted substrate heating for crack mitigation in laser metal deposition”, Applied Physics. A, Materials Science & Processing, v. 131, n. 12, pp. 991, 2025. doi: https://doi.org/10.1007/s00339-025-09122-5.
  » https://doi.org/10.1007/s00339-025-09122-5
• ^[4] WAN, X., LI, J., ZHANG, J., et al, “Effect of the process parameters on deposition quality of Ti6Al4V titanium alloy by powder-feed underwater laser metal deposition”, Welding in the World, v. 69, n. 11, pp. 3455–3465, 2025. doi: https://doi.org/10.1007/s40194-025-02160-w.
  » https://doi.org/10.1007/s40194-025-02160-w
• ^[5] ZHOU, J., SONG, Y., “Ultrafast laser asynchronous processing for laser metal deposition manufacturing of Ti6Al4V”, International Journal of Advanced Manufacturing Technology, v. 138, n. 9–10, pp. 4675–4688, 2025. doi: https://doi.org/10.1007/s00170-025-15819-x.
  » https://doi.org/10.1007/s00170-025-15819-x
• ^[6] NOVAK, M., CELBA, M., CERMAK, A., et al, “Quality of weld bead in laser metal deposition of titanium wire using pulsed laser”, Progress in Additive Manufacturing, v. 10, n. 8, pp. 4705–4720, 2025. doi: https://doi.org/10.1007/s40964-024-00855-y.
  » https://doi.org/10.1007/s40964-024-00855-y
• ^[7] DONG, F., KONG, L., WANG, H., et al, “Cross-section geometry prediction for laser metal deposition layer-based on multi-mode convolutional neural network and multi-sensor data fusion”, Journal of Manufacturing Processes, v. 108, pp. 791–803, 2023. doi: https://doi.org/10.1016/j.jmapro.2023.11.036.
  » https://doi.org/10.1016/j.jmapro.2023.11.036
• ^[8] WU, D., LUO, G., HU, Y., et al, “Printing regime for single metal microdroplet deposition in laser-induced forward transfer.”, Optics and Lasers in Engineering, v. 167, pp. 107617, 2023. doi: https://doi.org/10.1016/j.optlaseng.2023.107617.
  » https://doi.org/10.1016/j.optlaseng.2023.107617
• ^[9] CABREIRA, V., SANTANA, R.M.C., “Effect of infill pattern in Fused Filament Fabrication (FFF) 3D Printing on materials performance”, Matéria (Rio de Janeiro), v. 25, n. 3, pp. e12826, 2020. doi: https://doi.org/10.1590/s1517-707620200003.1126.
  » https://doi.org/10.1590/s1517-707620200003.1126
• ^[10] ABBAS, A.H., MORTAZAVI, S.Z., REYHANI, A., et al, “Engineering WS2/MoS2 thin films via pulsed laser deposition: structural and optical characterization”, Journal of Materials Science Materials in Electronics, v. 36, n. 18, pp. 1128, 2025. doi: https://doi.org/10.1007/s10854-025-15189-w.
  » https://doi.org/10.1007/s10854-025-15189-w
• ^[11] WANG, Z., CHEN, M., ZHAO, K., et al, “Effect of different feedstocks on the microstructure and mechanical properties of HSLA steel repaired by underwater laser direct metal deposition”, Materials Chemistry and Physics, v. 314, pp. 128935, 2024. doi: https://doi.org/10.1016/j.matchemphys.2024.128935.
  » https://doi.org/10.1016/j.matchemphys.2024.128935
• ^[12] SAZHIN, A., DUBROV, A., OZHEREDOV, I., “Optical emission spectroscopy for in-line control of laser additive manufacturing with metal wire of titanium-nickelide and magnesium alloys”, Optical and Quantum Electronics, v. 57, n. 11, pp. 582, 2025. doi: https://doi.org/10.1007/s11082-025-08507-3.
  » https://doi.org/10.1007/s11082-025-08507-3
• ^[13] GAO, L., ZHANG, H., ZHI, H., et al, “Corrosion characteristics and corrosion depth predicted of rolled/laser metal deposition additive manufactured Ti-6Al-4V joints”, Materials Today. Communications, v. 44, pp. 111897, 2025. doi: https://doi.org/10.1016/j.mtcomm.2025.111897.
  » https://doi.org/10.1016/j.mtcomm.2025.111897
• ^[14] SILVA, G.C.D., PEIXOTO, G.C., DURÃES, G.G.S., et al, “Comparative mechanical analysis of polymers made by SLA, FDM, and injection”, Matéria (Rio de Janeiro), v. 31, pp. e20250817, 2026. doi: https://doi.org/10.1590/1517-7076-rmat-2025-0817.
  » https://doi.org/10.1590/1517-7076-rmat-2025-0817
• ^[15] XU, W., DU, Z., WANG, Z., et al, “A comparative study on the corrosion, wear, and tribocorrosion performances of Ti-6.5Al-2Zr-1Mo-1V (TA15) alloy fabricated by selective laser melting and laser metal deposition”, Journal of Materials Research and Technology, v. 30, pp. 6245-6253, 2024. doi: https://doi.org/10.1016/j.jmrt.2024.05.049.
  » https://doi.org/10.1016/j.jmrt.2024.05.049
• ^[16] LEKOADI, P.M., TLOTLENG, M., SIYASIYA, C.W., et al, “Investigation of network microstructure of TiB/Ti6Al4V-ELI composite manufactured with laser metal deposition”, MRS Advances, v. 8, n. 10, pp. 602–606, 2023. doi: https://doi.org/10.1557/s43580-023-00576-4.
  » https://doi.org/10.1557/s43580-023-00576-4
• ^[17] BORODKOV, A.S., SIMAKINA, Y.I., GRECHNIKOV, A.A., “Surface-assisted laser desorption/ionization of metal complexes with dithizone”, Journal of Analytical Chemistry, v. 79, n. 11, pp. 1574–1578, 2024. doi: https://doi.org/10.1134/S1061934824700989.
  » https://doi.org/10.1134/S1061934824700989
• ^[18] ZHANG, Q., WANG, B., ZHAO, M., et al, “A material removal model considering the effect of anisotropy on the processing of laser metal deposition Ti-alloy”, International Journal of Advanced Manufacturing Technology, v. 127, n. 5–6, pp. 2785–2797, 2023. doi: https://doi.org/10.1007/s00170-023-11654-0.
  » https://doi.org/10.1007/s00170-023-11654-0
• ^[19] LI, W., ZHENG, K., ZHAO, G., et al, “Variant selection in the thermally treated Ti-6Al-4V processed by laser metal deposition”, Metallography, Microstructure, and Analysis, v. 12, n. 4, pp. 645–650, 2023. doi: https://doi.org/10.1007/s13632-023-00975-x.
  » https://doi.org/10.1007/s13632-023-00975-x
• ^[20] CHINNATHAMBI, D., JAGANATHAN, S., RAJENDRAN, S., et al, “Investigation of material characteristics on 3D printing of vertical windmill blade using finite element method”, Matéria (Rio de Janeiro), v. 29, n. 4, pp. e20240365, 2024. doi: https://doi.org/10.1590/1517-7076-rmat-2024-0365.
  » https://doi.org/10.1590/1517-7076-rmat-2024-0365
• ^[21] WEBER, J.U., EMMELMANN, C., “Localized defect frequencies for Laser Metal Deposition processes”, Procedia CIRP, v. 124, pp. 331–334, 2024. doi: https://doi.org/10.1016/j.procir.2024.08.128.
  » https://doi.org/10.1016/j.procir.2024.08.128
• ^[22] PACQUENTIN, W., WIDENT, P., VARLET, J., et al, “Temperature influence on the repair of a hardfacing coating using laser metal deposition and assessment of the repair innocuity”, Journal of Advanced Joining Processes, v. 11, pp. 100284, 2025. doi: https://doi.org/10.1016/j.jajp.2025.100284.
  » https://doi.org/10.1016/j.jajp.2025.100284
• ^[23] SU, G., SHI, Y., LI, G., et al, “Improving the deposition efficiency and mechanical properties of additive manufactured Inconel 625 through hot wire laser metal deposition”, Journal of Materials Processing Technology, v. 322, pp. 118175, 2023. doi: https://doi.org/10.1016/j.jmatprotec.2023.118175.
  » https://doi.org/10.1016/j.jmatprotec.2023.118175
• ^[24] BOAVENTURA, L.B., REINKE, G., OLIVEIRA, A.C.A.D., et al, “Influence of support structures parameters on surface quality in FDM parts”, Matéria (Rio de Janeiro), v. 30, pp. e20250688, 2025. doi: https://doi.org/10.1590/1517-7076-rmat-2025-0688.
  » https://doi.org/10.1590/1517-7076-rmat-2025-0688
• ^[25] AL-MAHMOOD, M., SAKHI, S., ABUZAID, W., et al, “Synergistic integration of carbon nanotubes in laser powder bed fusion: advancing multifunctional metal additive manufacturing”, Journal of Materials Science, v. 60, n. 33, pp. 14279-14338, 2025. doi: https://doi.org/10.1007/s10853-025-11175-z.
  » https://doi.org/10.1007/s10853-025-11175-z
• ^[26] LOKESH, P.C., BALDACI, G., DU, X., et al, “A parametric model to optimize ultrafast laser fabrication of hydrophobic metal surfaces”, Lasers in Manufacturing and Materials Processing, v. 12, n. 2, pp. 369–384, 2025. doi: https://doi.org/10.1007/s40516-025-00290-z.
  » https://doi.org/10.1007/s40516-025-00290-z
• ^[27] SADHU, A., RAMADAS, H., PANDA, B.K., et al, “Laser surface polishing of additively manufactured NiCrSiBC-60WC ceramic-metal composite using a line beam”, Journal of Materials Engineering and Performance, v. 34, n. 22, pp. 27231–27245, 2025. doi: https://doi.org/10.1007/s11665-025-11289-0.
  » https://doi.org/10.1007/s11665-025-11289-0
• ^[28] MATVEEV, A.E., PROMAKHOV, V.V., SCHULZ, N.A., et al, “Structure and phase composition of materials fabricated by direct laser deposition using Al-TiB2 composite powder”, Russian Physics Journal, v. 68, n. 3, pp. 400–407, 2025. doi: https://doi.org/10.1007/s11182-025-03446-8.
  » https://doi.org/10.1007/s11182-025-03446-8
• ^[29] PAN, W., XU, B., LI, C., “Effects of groove shape on microstructure and mechanical responses of laser-directed energy deposition-repaired GH4099 Ni-Based superalloy”, Acta Metallurgica Sinica, v. 38, n. 6, pp. 1003–1011, 2025. doi: https://doi.org/10.1007/s40195-025-01837-x.
  » https://doi.org/10.1007/s40195-025-01837-x
• ^[30] ABADI, S.N.R., MI, Y., KISIELEWICZ, A., et al, “Influence of laser-wire interaction on heat and metal transfer in directed energy deposition”, International Journal of Heat and Mass Transfer, v. 205, pp. 123894, 2023. doi: https://doi.org/10.1016/j.ijheatmasstransfer.2023.123894.
  » https://doi.org/10.1016/j.ijheatmasstransfer.2023.123894
• ^[31] MANI, N., TRAN, N., JONES, A., et al, “ Additive manufacturing of titanium-diamond parts: insights into the laser metal deposition process, powder rheology, mechanical properties and osteoblast cell viability”, Rapid Prototyping Journal, v. 30, n. 10, pp. 1989–2006, 2024. doi: https://doi.org/10.1108/RPJ-10-2023-0357.
  » https://doi.org/10.1108/RPJ-10-2023-0357
• ^[32] MAFFIA, S., CHIAPPINI, F., MAGGIANI, G., et al, “Enhancing productivity and efficiency in conventional laser metal deposition process for Inconel 718 - Part II: advancing the process performance”, International Journal of Advanced Manufacturing Technology, v. 129, n. 1–2, pp. 279–298, 2023. doi: https://doi.org/10.1007/s00170-023-12197-0.
  » https://doi.org/10.1007/s00170-023-12197-0
• ^[33] SU, G., SHI, Y., LI, G., et al, “Highly-efficient additive manufacturing of Inconel 625 thin wall using hot-wire laser metal deposition: process optimization, microstructure, and mechanical properties”, Optics & Laser Technology, v. 175, pp. 110763, 2024. doi: https://doi.org/10.1016/j.optlastec.2024.110763.
  » https://doi.org/10.1016/j.optlastec.2024.110763
• ^[34] ILANLOU, M., RAZAVI, R.S., HAGHIGHAT, S., et al, “Multi-track laser metal deposition of Stellite6 on martensitic stainless steel: geometry optimization and defects suppression”, Journal of Manufacturing Processes, v. 86, pp. 177–186, 2023. doi: https://doi.org/10.1016/j.jmapro.2022.12.036.
  » https://doi.org/10.1016/j.jmapro.2022.12.036
• ^[35] SINGH, A.K., BAL, K.S., DEY, D., et al, “Experimental investigation and parametric optimization for minimization of dilution during direct laser metal deposition of tungsten carbide and cobalt powder mixture on SS304 substrate”, Powder Technology, v. 390, pp. 339–353, 2021. doi: https://doi.org/10.1016/j.powtec.2021.05.056.
  » https://doi.org/10.1016/j.powtec.2021.05.056
• ^[36] MORADI, A., TAJALLI, S., MOSALLANEJAD, M.H., et al, “Intelligent laser-based metal additive manufacturing: A review on machine learning for process optimization and property prediction”, International Journal of Advanced Manufacturing Technology, v. 136, n. 2, pp. 527–560, 2025. doi: https://doi.org/10.1007/s00170-024-14858-0.
  » https://doi.org/10.1007/s00170-024-14858-0
• ^[37] PAGAN, M., HUTCHINSON, A., LIES, N.J., et al, “Improvement of powder blown laser beam-directed energy deposition of molybdenum using lanthanum oxide additions”, Journal of the Minerals Metals & Materials Society, v. 77, n. 10, pp. 7235–7246, 2025. doi: https://doi.org/10.1007/s11837-025-07343-w.
  » https://doi.org/10.1007/s11837-025-07343-w
• ^[38] BALLA, S.K., MALLAIAH, M., NAGAMUTHU, S., et al, “The influence of laser direct energy deposition processing parameters on Al7075 alloy and Zr-modified Al7075 alloy”, International Journal of Advanced Manufacturing Technology, v. 135, n. 1–2, pp. 181–201, 2024. doi: https://doi.org/10.1007/s00170-024-14503-w.
  » https://doi.org/10.1007/s00170-024-14503-w
• ^[39] ZHONG, C., REN, X., FU, K., et al, “Effect of thermal aging on the microstructure and mechanical properties of Inconel 625 superalloy fabricated by high-speed laser metal deposition”, Metallurgical and Materials Transactions. B, Process Metallurgy and Materials Processing Science, v. 54, n. 5, pp. 2723–2736, 2023. doi: https://doi.org/10.1007/s11663-023-02869-3.
  » https://doi.org/10.1007/s11663-023-02869-3
• ^[40] HECZEL, A., BARTON, D.J., GUBICZA, J., et al, “Evolution of defects and microstructure in AISI 4340 steel produced by laser-assisted cold spray deposition”, Journal of Thermal Spray Technology, v. 35, n. 1, pp. 281–296, 2026. doi: https://doi.org/10.1007/s11666-025-02113-1.
  » https://doi.org/10.1007/s11666-025-02113-1
• ^[41] REZAEE HAJIDEH, M., FARAHANI, M., “Evaluating the effect of repair using direct laser metal deposition on the fracture resistance of the cracked 1.2714 tool steel component”, Journal of the Brazilian Society of Mechanical Sciences and Engineering, v. 46, n. 3, pp. 125, 2024. doi: https://doi.org/10.1007/s40430-024-04686-4.
  » https://doi.org/10.1007/s40430-024-04686-4
• ^[42] ZHOU, Z., GU, J., TONG, D., “Numerical simulation of residual stress evolution and interfacial cracking in directed laser deposition of γ-TiAl alloys”, Metallurgical and Materials Transactions. A, Physical Metallurgy and Materials Science, v. 56, n. 12, pp. 5683-5694, 2025. doi: https://doi.org/10.1007/s11661-025-08006-w.
  » https://doi.org/10.1007/s11661-025-08006-w
• ^[43] LUO, G., WU, D., ZHOU, Y., et al, “Alternate deposition and remelting microdroplets via single laser for printing low-defect and high-performance metal micropillars”, International Journal of Machine Tools & Manufacture, v. 197, pp. 104136, 2024. doi: https://doi.org/10.1016/j.ijmachtools.2024.104136.
  » https://doi.org/10.1016/j.ijmachtools.2024.104136

Corresponding Author:

Sivaramakrishnan Arumugam, [e-mail], [e-mail], [e-mail], [e-mail]

Publication Dates

• Publication in this collection
  18 Sept 2026
• Date of issue
  2026

History

• Received
  18 Mar 2026
• Accepted
  03 Aug 2026