1

Parametric optimisation and microstructural analysis on high power Yb-fibre laser welding

of Ti-6Al-4V

J. Ahn1*, L. Chen2, C. M. Davies1 and J. P. Dear1

1 Department of Mechanical Engineering, Imperial College London, South Kensington Campus,

London, UK SW7 2AZ

2 Science and Technology on Power Beam Lab, Beijing Aeronautical Manufacturing Technology

Research Institute, China

Abstract

In this work thin sheets of Ti-6Al-4V were full penetration welded using a 5 kW fibre laser in order to

evaluate the effectiveness of high power fibre laser as a welding processing tool for welding Ti-6Al-4V

with the requirements of the aircraft industry and to determine the effect of welding parameters

including laser power, welding speed and beam focal position on the weld microstructure, bead profile

and weld quality. It involved establishing an understanding of the influence of welding parameters on

microstructural change, welding defects, and the characteristics of heat affected zone (HAZ) and weld

metal (WM) of fibre laser welded joints. The optimum range of welding parameters which produced

welds without cracking and porosity were identified. The influence of the welding parameters on the

weld joint heterogeneity was characterised by conducting detailed microstructural analysis.

Key words

Ti-6Al-4V; Fibre laser; Welding; Microstructure; Titanium Alloy; Porosity

*corresponding author: joseph.ahn08@imperial.ac.uk; Tel: 020 7594 7035; Fax: 020 7594 7017

2

1. Introduction

The quality and performance of welded joints depend on the weld geometry, melt pool behaviour

during welding, the metallurgy of the welded zone and the heat affected zone, and welding defects.

The complexity of chemical and metallurgical actions which take place during welding may result in

subsequent failure of the weld in service and so, it is important to anticipate and incorporate the

effects of welding at the design stage. Around 45% of the causes for welding imperfections is due to

poor process conditions (1). In fact, because the research on the development of welding techniques

has been largely conducted to satisfy the needs of the industry for demonstrating the maximum

capabilities of a process, many welding fundamentals have not been researched and therefore are

not yet fully understood (2). This means that a significant amount of work is required to be done to be

able to predict and optimise the laser welding process to produce consistently quality welds.

In order to obtain an acceptable weld profile and satisfactory mechanical properties, control of weld

bead shape is essential as the mechanical properties of welds are affected by the weld bead shape.

The weld bead shape which affects the weld metal solidification behaviour is influenced by welding

parameters and the corresponding amount of heat input into the workpiece (3). Therefore, it is

necessary to determine the influence of the welding parameters including laser power, welding speed

and defocusing distance (4,5) on weld morphology as well as to identify the sources of welding

defects. It would then be possible to identify the optimum combination of welding parameters which

ensures the required weld quality and properties, and also minimises welding defects (6–8).

This investigation was concerned with the welding parameters including laser power, welding speed

and focal distance, and their effects on the weld shape and the final solidification structures of Ti-6Al-

4V welds which further influence the overall mechanical properties of the welds. Integrated parametric

studies which focus on the influence of various welding parameters on the welding results for fibre

laser welding of Ti-6Al-4V have been rarely published. In many cases, even for conventional lasers,

investigations have often been conducted case specific with a certain range of process parameters,

thus, making it difficult to transfer the results.

Previous studies concerning the effect of welding parameters on weld seam geometry have identified

different shapes of laser welds in terms of the top and bottom weld widths. The work of Karlsson et al.

(9–13) for example, investigated the influence of laser welding parameters on the weld seam

3

geometry of high strength steels welded with 15 kW fibre laser and it was found that increasing the

welding speed can suppress root sagging and undercuts while decreasing the welding speed can

suppress lack of penetration. Manonmani et al. (14) studied the effect of laser power, welding speed

and beam incidence angle on the weld seam geometry in terms of the penetration depth, the bead

width and the area of penetration of a 2.5 mm thick AIS1304 stainless steel. They showed that the

depth of penetration and penetration area increased with the laser power and the beam angle,

whereas, the weld width decreased with increasing welding speed. A trend was observed where the

penetration depth and area increased to a maximum value and then decreased with any further

increase in welding speed. This was due to the fact that the mode of heat transfer changes from a

keyhole mode at lower speeds to a conduction mode at higher speeds.

Different top and root shape classes of laser welded Ti-6Al-4V have been identified in published

literature. The centre of the fusion zone as observed by Balasubramaninan et al. (15) in CO2 laser

welded Ti-6Al-4V joints, generally presented a convex shape at middle thickness due to volume

contraction, surface tension and phase transformation during welding. The specific heat input as

mentioned above has a strong impact on the welding bead shape. An increase in welding speed, due

to the lower value of specific heat input transmitted to the workpiece, leads to a reduction of the weld

width. The heat input was found to be a highly influencing parameter for the bead shape by Squillace

et al., (16) who studied the effect of welding parameters on morphology and mechanical properties of

fibre laser welded Ti-6Al-4V butt joints. Higher heat input promoted an hour glass shaped weld bead

whereas, lower heat input promoted the formation of a V-shaped weld bead. Campanelli et al.

investigated fibre laser welding of Ti-6Al-4V at a constant power of 1.2 kW and they observed a

change in the weld shape from nail head to V-shape when the welding speed was increased from 2 to

2.5 m/min (22). Mueller et al. (18,19) studied the potential application of CO2 laser and fibre laser

welding of Ti-6Al-4V for aircraft applications and observed a trend where a change in the weld shape

from an hourglass shape at low speed to nail head shape at fast speed increase the tendency to

entrap gases and form root porosity. The same trend was also observed by Chen et al. (7) in their

investigation on CO2 welding BT20 titanium alloy. Hilton et al. (20) investigated fibre laser welding of 3

and 5 mm thick Ti-6Al-4V and linked the weld profile to porosity level. Low levels of porosity were

found in the 5 mm thick weld which used high laser powers. Interestingly, it had a face weld width

smaller than the root weld width due to the keyhole behaviour which produced a larger molten volume

4

in the lower part of the weld and promoted the escape of any trapped root shielding gas (20). The

investigation presented in this paper was therefore conducted to determine the influence of key fibre

laser welding parameters on the resultant weld quality when welding Ti-6Al-4V, the formation of

welding defects and to develop techniques to avoid these issues.

2. Materials and Experimental Procedures

2.1 Materials

Mill-annealed titanium alloy Ti-6Al-4V (Grade 5) sheets with 2 mm thickness were used with nominal

composition given in Table 1. Ti-6Al-4V welds were produced under various welding conditions of

laser power, welding speed and focal position. A wide range of processing parameters were tested as

illustrated in Table 4, Table 5 and Table 6 in order to maximise the operating window and produce a

good weld quality by identifying the optimum combination of laser power, welding speed and focal

position. All Ti-6Al-4V welding trials were autogenous.

A 5 kW continuous wave (CW) ytterbium fibre laser was used in TEM01* mode for laser welding as

shown in Figure 1. The beam diameter at focus was 630 μm, the wavelength of fibre laser was

1070 nm, the beam quality factor, M2 was around 7.3, the divergence half angle of the focused beam

was 12.5 mrad and the Rayleigh length for the multimode beam, scaled with the M2 factor was around

22.6 mm. A beam parameter product (BPP) of less than 2.5 mm mrad was formed. All these

parameters influence the focusability and absorptivity of the laser beam.

Industrial grade argon with 99.999% purity was used for welding Ti-6Al-4V. The shielding gas was

supplied to protect both top and underside of the weld. The coaxial shielding gas was delivered via

the weld nozzle to protect molten pool and the back protecting shield gas was supplied via the

shielding gas path in the copper insert to protect back weld. A total shielding of titanium weld seam

was provided by the additional trailing shielding shoe with its own shielding gas supply following along

behind the laser beam at a distance no more than 2 mm from the surface of the workpiece (22).

Argon was used for back shielding in all experiments at a flow rate of 15-20 l/min. The flow rate for

coaxial or side-jet shielding was 20-25 l/min and 25 l/min for drag cover shielding.

Welding was performed parallel to the rolling direction of the Ti-6Al-4V sheets, and followed the

recommendations written in the standards ISO/TR 17671-6, AWS C7.2 and BS EN 1011-6 for the

5

control of laser beam welding of metallic materials (23–25). All the specimens which were used for

microstructural analysis were welded bead on plate so there was no requirement for joint fit up or

alignment tolerances during weld preparations. The main welding parameters investigated were laser

power, welding speed and focal distance.

2.2 Metallographic Specimen Preparation

Preparation of weld specimens for metallographic examinations was conducted by following the

methods specified in the standards, ASTM E3 (26) and BS EN 1321 (27). Two transverse cross-

sections were taken from the welds by electrical discharge machining (EDM) for all welding

parameters set, one for examining the weld seam top surface and the other for examining the weld

cross-section. The polished specimens were chemically etched for optical microscopy. The

microstructural constituents of the weld were revealed by using suitable chemical etchants. The

recommended procedure was to use Kroll’s reagent which is a mixture of 92% distilled water, 6%

HNO3, 2% HF for etching Ti-6Al-4V specimens by dipping for 15 seconds (28). However, such

duration was found to be suitable for etching the parent material but not long enough for the weld

metal. For this reason, etching was performed for a longer duration in the range of 30 to 60 s

depending on the specimen until the true metallographic weld microstructure of the weld was

sufficiently revealed for microstructural analysis.

2.3 Experimental Procedures

Experiments were conducted in accordance with the requirements specified in BS EN ISO 15614-11

and BS EN ISO 13919-2 (29-30). These standards provide guidance on levels of imperfections in

electron and laser beam welded joints in metallic materials. The relevant examination and tests for

welds in accordance with acceptance level D are listed in Table 2.

As Table 2 shows, visual examination on the welded sheets and metallographic examination on

welded specimens were conducted. Since a minimum of one section extracted from the weld is

required for metallographic examination, it was therefore necessary to check the conformity of the

weld seams. A single weld cross section per welding parameter set cannot guarantee the uniformity

of weld seam geometry over the whole weld seam. In order to obtain representative trends in

variations of the weld seam geometry for varying welding parameters, the conformity of weld seam

geometry must be checked. The possible methods of checking the conformity of weld seam geometry

6

over the whole weld same include extracting sections parallel to the welding direction or making a

large number of samples with the same parameters. However, since a large number of parameter

variations were investigated in this study, it was not practical to follow these methods which consume

a lot of time, effort and material costs. Instead, the shape and the quality of the top and bottom weld

seam surfaces were used as an indicator for the conformity of the weld seam geometry. Random

inspections of weld seams showed a positive relationship between the quality of the top and bottom

weld seam surfaces and the uniformity of the weld geometry. Three cross sections were extracted;

one at the centre, one near the weld start and the other close to the weld end positions. The shape of

the weld seams remained fairly constant and only small variations were observed on the top and

bottom weld seam surfaces. Therefore, the shape of the top and bottom weld seam surfaces were

used to evaluate the conformity of the whole weld, making it possible to assume that the chosen weld

cross section is representative for the whole weld.

Macroscopic and microscopic inspection of transverse sections of metallographic welded specimens

followed the test procedures specified in BS EN 1321 (27). Macroscopic examination was conducted

under low magnification with etching and microscopic examination with a magnification within 50 to

500 times also with etching to reveal features of welded joints. An optical microscope (OM) was used

for microstructural examination and an environmental scanning electron microscope (SEM) with

energy dispersive X-ray spectroscopy (EDX) was used to determine chemical compositions of the

welded specimens at an accelerating voltage of 15 kV, an emission current of 76 μA, a working

distance of 6.8 mm, an elevation of 35° and a live time of 50 seconds.

2.4 Welding Quality Acceptance Criteria

The quality of welds produced in this experiment were evaluated against a set of welding acceptance

criteria from several international standards on welding as illustrated in Table 3. The American

Welding Society standard, AWS D17.1:2001 (31) was applied to laser beam welding of Ti-6Al-4V,

however, the standard was developed for general purpose fusion welding and not specifically for laser

beam welding. The European Standard, BS EN ISO 13919-2:2001 (30) was developed specifically for

electron beam and laser beam welding of aluminium based alloys but not for titanium based alloys.

Another European Standard BS EN 4678:2011 was developed specifically for laser beam welding of

metallic materials for aircraft applications and it includes welding quality acceptance criteria for

7

titanium based alloys. It is often the case where for aerospace applications, quality acceptance criteria

more stringent than those cited by these standards are required and therefore, the most stringent

criteria from each standard were used (20).

In all three standards, welds containing cracks, incomplete penetration or incomplete fusion are

rejected without tolerances. Porosity and undercut especially, can act as stress concentrations so

limits are placed in the standards on both weld porosity levels and undercut depth. Undercutting is a

common defect in laser beam welding due to its high processing speed which leads to insufficient

time for the molten base metal to flow back to the weld toe and fill the gap. Undercut, underfill, root

concavity and shrinkage grooves are particularly undesirable for components subject to dynamic

loading as these defects act as stress concentrators and subsequently initiate fatigue cracks.

Welding quality acceptance criteria on weld face width and root width are only specified in AWS D.17.

In addition to this, the ratio of root width to face width (Rw) can also be used to evaluate the

processing stability of full penetration welding. An investigation conducted by Chen et al. on Nd:YAG

laser welding on 3 mm thick 5A90 Al-Li alloy and 2.5 mm thick BT20 titanium alloy proved that a weld

width ratio of greater than 0.4 for the titanium alloy produced the conditions for a stable full

penetration welding. (7). Therefore, the recommended value of 0.4 was used to evaluate the weld

shapes of Ti-6Al-4V.

3. Results and Discussion on Ti-6Al-4V

In all the Ti-6Al-4V welded specimens investigated in this experiment, weld underfill and root

concavity were not present so every examined specimen passed the weld quality assessment criteria

on underfill and root concavity. On the other hand, excess weld metal (top) or penetration (root) was

observed in all the specimens to a certain degree, but still much below the limits of 0.66, 0.50 and

0.40 mm defined in Table 3. The level of porosity detected was less than the size limit set for porosity

of 0.66, 0.60 and 0.40 mm in Table 3 for all specimens. The largest pore observed had a diameter of

0.23 mm, which still met the requirements of Table 3. It was found that if any porosity was evident, it

occurred in the lower half of the weld. The way in which this can be reduced is by improving the

shielding gas arrangements especially the under-bead shielding gas.

No cracking defect was observed in all Ti-6Al-4V welded specimens as expected because titanium

alloys are generally not considered susceptible to fusion zone solidification cracking and HAZ

8

liquation cracking due to the absence of secondary phase dispersoids or precipitate particles, or

impurities at the grain boundaries (32). Of more concern was embrittlement at high elevated

temperatures arising from contamination by hydrogen, nitrogen and oxygen absorption. An effective

trailing shield was used to prevent this problem. Therefore, in this investigation the weld joint

discontinuities were assessed mainly in terms of undercut, porosity, reinforcement, face and root

widths and the ratio of root to face width (Rw).

Figure 2 shows the microstructures of base metal (BM), heat affected zone (HAZ) and fusion zone

(FZ) obtained from a Ti-6Al-4V weld cross-section produced using a laser power of 2.7 kW, welding

speed of 2.1 m/min and a focal position of +4 mm. The microstructure of the as-received mill-

annealed Ti-6Al-4V at equilibrium and in room temperature consisted of mainly equiaxed α grains with

some retained intergranular β phase as shown in Figure 2, where the bright regions are the equiaxed

α and the dark regions are the intergranular β distributed at the elongated α grain boundaries. The

equiaxed microstructure was obtained by mechanical working in the α-β phase field to transform

lamellar α into equiaxed α in a transformed β matrix. The vanadium enriched β-phase was only

present in limited quantity at room temperature meaning that it made only a small contribution

strengthening because of its small proportion.

The microstructure of the HAZ between the FZ and the BM consisted of two sub-regions, one close to

the FZ and the other close to the BM. The HAZ close to the FZ experienced higher temperatures

during welding ranging from the β transus temperature of 980-995°C to the solidus temperature of

1605°C (33,34), whereas, the HAZ close to the BM was subjected temperatures below the β transus

temperature but high enough to affect its microstructure. The observed HAZ near the FZ consisted

mostly of acicular martensitic phase α´ and a small amount of acicular α and primary α similar to the

FZ. Wang et al. (35) suggested from their results on CO2 laser welding of 3.3 mm thick Ti-6Al-4V with

a laser power of 2.5 kW and welding speed of 1.5 m/min, that this kind of microstructure corresponds

to a specimen quenched from a temperature below the beta transus, similar to that observed by water

quenching from 1100°C. However, it was difficult to determine precisely the range of temperatures

from which the cooling has occurred. The microstructure of the HAZ close to the BM consisted of a

mixture of the microstructural constituents in the FZ, a small amount of martensite α´, and the BM,

mainly primary α and intergranular β phase (36). The volume fraction of α´ in the HAZ decreased with

increasing distance from the FZ boundary which was in agreement with the results from the work by

9

Kabir et al. (37) on Nd:YAG laser welding of 3.2 mm thick Ti-6Al-4V. The microstructure of the FZ

contained fine acicular α´ with coarse columnar prior β grain boundaries which originated during weld

solidification and grew opposite to the direction of heat flow from the partially melted β grains in the

near HAZ to the weld centreline as shown in Figure 2. The absence of grain boundary phase α along

the prior β grain boundaries in the FZ according to Kabir et al. (38) indicates that the cooling rate after

welding was greater than the critical cooling rate for diffusional transformation of 410°C/s in Ti-6Al-4V

as determined by Ahmed et al. (39). Kabir et al. (38) confirmed the presence of grain boundary α

along the prior β grains in Nd:YAG laser welded 5.1 mm thick Ti-6Al-4V at a laser power of 4 kW and

welding speed of 1 m/min because the cooling rate was close to the minimum limit necessary for

diffusionless transformation. High cooling rates from temperatures above the martensite start

temperature associated with the laser beam welding process promotes transformation of the body

centre cubic (BCC), β into the hexagonal close packed (HCP) α´ microstructure.by a diffusionless

transformation process (28). The α´ is a metastable supersaturated α phase with an acicular

morphology and its volume fraction increases with increasing cooling rate. Due to its formation by

rapid nucleation and growth, it contains a higher dislocation density compared to the primary α grains

with a smaller plate thickness. The smaller grain size and the higher dislocation density of α´ are

responsible for the higher hardness of α´ than α (40).

Energy-dispersive X-ray spectroscopy was used for the elemental analysis of welded specimens.

Figure 3 shows the EDS spectrum of the fusion zone obtained from the scanning electron

microscopes (SEM) and the analysed chemical composition of the parent material, the heat affected

zone and the fusion zone. Minor differences in titanium, aluminium and vanadium contents were

found in these three zones. Even though small in difference, the percentage fraction of aluminium was

in a decreasing order from the PM to the FZ. Aluminium has the lowest boiling point of 2519°C, while

that of vanadium is 3407°C and titanium is 3287°C. As the maximum temperature during welding can

reach above the boiling point of the material, it was determined that the loss of alloying elements by

vaporisation was dependent on the boiling point of each element. As a result, the fraction of

vanadium, with the highest boiling point, increased towards the weld, whereas, the other two

decreased according to their vaporisation temperature.

10

3.1 Effect of Laser Power

Various laser powers were employed as listed in Table 4 to examine the effect of laser power on the

weld quality and microstructure. Figure 4 shows the effect of changing laser power at a low welding

speed of 2.1 m/min on the weld transverse sections. Different laser powers resulted in different weld

shapes. The two horizontal lines represent the requirements set by the welding quality acceptance

criteria in BS EN 4678 in Table 3 for the face width and the root width. The face width is required to

be below the blue line, and the root width in between the red and blue lines. The only laser power

which failed these criteria was 1.2 kW. Unlike all the laser powers, where full penetration was

obtained, incomplete penetration was observed at 1.2 kW. Incomplete penetration resulted from

insufficient heat input supplied to the workpiece at 1.2 kW. Both the top and bottom weld widths

increased with laser power up to 2.4 kW. and then decreased at higher laser powers. The weld shape

changed from a V shape at lower laser powers to an hourglass shape at higher laser powers, as a

result of increased heat input transmitted to the material as shown in Figure 5. The trend in general

was in a reasonable agreement with the findings of Squillace et al. (16), Campanelli et al. (17), and

Mueller et al. (18,19).

The Rw values as shown in in Figure 4 were above 0.4 and the welding process was stable at all laser

powers. except at 1.2 kW due to incomplete penetration. In contrast, undercut was observed at all

laser powers, except again at 1.2 kW. The depth of undercut increased with increasing laser power up

to a maximum of around 0.15 mm. The criteria for undercut as indicated by the two red horizontal

lines, where the lower one corresponds to the 0.05 mm limit specified in AWS D17.1 at 0.05 mm and

the upper one corresponds to the 0.10 mm limit specified in BS EN ISO 13919-2 and BS EN

4678.Undercut produced in at laser powers above 1.5 kW were all greater than the maximum limit

specified in AWS D17.1, but acceptable accepted according to BS EN ISO 13919-2 and BS EN 4678

at laser powers less than or equal to 1.8 kW This meant that the optimum processing parameter

window was very small for the given constant slow welding speed of 2.1 m/min since the depth of

undercut was only acceptable at 1.5 kW or 1.8 kW. The main cause of increased depth of undercut

was excessive power or heat input, which enhanced evaporation and expulsion of molten materials

from the sides of the welded joint into the weld centre, leaving a drain-like impression along the length

of the weld (35,37,41). The influence of laser power on weld reinforcement was not significant at all

laser powers.

11

Porosity was observed at 1.2, 1.5 and 1.8 kW as shown in Figure 5, in the specimens which showed

either lack of penetration or relatively narrow root width. It is well known that the welds without full

penetration are more likely to contain a higher amount of porosity (42). For example, according to

Murav’ev (43), pores account for 43-56% of the total number of defects. The formation of porosity was

found to be depended on heat input. Increasing the laser power also increased the heat input, which

effectively reduced the solidification rate to allow more time for gases to escape from the weld pool

and thereby minimising the risk of porosity (44). The pores were found along the FZ centreline in the

lower half of the weld with a pore diameter of 0.10, 0.05 and 0.13 mm, respectively. The size of these

pores were smaller than the maximum tolerable size of 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. They were all spherical in shape and since there is

no element in Ti-6Al-4V with low boiling point, it was concluded that porosity was caused by the

presence of gases such as hydrogen, nitrogen, oxygen or argon shielding gas from the environment

and contamination of the workpiece. Hydrogen solubility in Ti-6Al-4V is a function of temperature and

as the temperature drops during solidification, the solubility of hydrogen increases, which promotes

the rejection of hydrogen into the FZ and pores are formed when the absorbed hydrogen cannot

escape the molten weld pool before solidification (45,46). The small size of pores measured indicated

that they were formed through diffusion of hydrogen (47). It is commonly thought that an hourglass

shaped cross-section has a greater tendency to entrap gases than nail head shaped welds due to the

downward sweeping solidification fronts in the root of hourglass shaped welds (7). However, in this

case, low internal porosity content was observed in hourglass shaped welds. On the other hand, the V

shaped welds produced at lower laser powers showed increasing propensity for porosity formation as

the escape of gas bubbles was only possible via the top surface and also because the solidification

rate increased as the weld width decreased, whereas, in fully penetrated welds including the

hourglass shaped ones, more effective escape was allowed via both the top and bottom surfaces

(44). Therefore, the best weld quality when welding at a constant welding speed of 2.1 m/min was

produced at a laser power of 1.8 kW.

Figure 6 shows the effect of changing laser power when a faster constant welding speed of 3.9 m/min

was used. The number of partially penetrated specimens increased when the welding speed

increased from 2.1 m/min to 3.9 m/min. These included the specimens welded at 1.8 and 2.1 kW,

which were previously fully penetrated at lower welding speeds but only partially penetrated at 3.9

12

m/min as a result of reduced heat input at a higher speed. The top width at all laser powers satisfied

the maximum limit of 3.2 mm specified in AWS D17.1 but the root width was smaller than the

minimum of 1.0 mm at laser powers less than or equal to 2.7 kW. The weld shape as shown in Figure

7 was narrow with a nail head shape at lower laser powers up to 3.0 kW and then wider with an

hourglass shape above 3.3 kW. The Rw increased from 0 to around 0.6 with increasing laser power up

to 3.0 kW, starting off with a nail head shape and then switching to an hourglass shaped weld. The Rw

at 1.8, 2.1 and 2.4 kW were less than 0.4, as proved by their nail head shape. Meanwhile, the depth

of undercut formed at these laser powers were satisfactory. The depth of undercut was below 0.10

mm at all laser powers. In addition, all welds confidently passed the criteria on reinforcement with no

clear trend with respect to laser power.

Porosity was only detected in the specimen welded at 2.1 kW as shown in Figure 7 with a diameter of

0.10 mm at the root of the weld, still considerably smaller than the most stringent size of 0.40 mm

specified in BS EN 4678. The cause of this porosity was identified to be due to hydrogen gas

entrapment. It was determined that a wider operating window was obtained when the welding speed

was increased to 3.9 m/min, passing all the criteria in BS EN ISO 13919-2 and BS EN 4678, and also

all criteria except that related to undercut in AWS D17.1 at laser powers above 3.0 kW.

The effect of changing laser power at a constant fast welding speed of 6.0 m/min is displayed in

Figure 8. Partial penetration was observed at 2.1 and 2.4 kW, whereas, previously at 2.1 m/min fully

penetrated hourglass shaped welds were produced and also at 3.9 m/min fully penetrated nail head

shaped weld was produced at 2.4 kW. The top weld width at all laser powers satisfied the maximum

limit of 3.2 mm, whereas, the bottom width at 2.1 to 3.0 kW were less than the minimum of 1.0 mm

required. The bottom width at above 3.4 kW were also very close to 1.0 mm. The V shape at 1.8 kW

changed to nail head shaped at 2.1 kW and remained the same up to 4.5 kW. In other words, a low

heat input transferred to the workpiece at all laser power levels. The Rw increased with laser power up

to 3.4 kW and then levelled off to a value close to 0.6 above 3.4 kW. The Rw at 2.1, 2.4 and 2.7 kW

were below 0.4, meaning that the process was unstable below 2.7 kW. Undercut at all laser powers

was very close to 0.05 mm limit specified in AWS D17.1. The laser power which satisfied all the

criteria on root width, Rw and undercut was 3.6 kW. Reinforcement measured at 3.6 kW was the

largest but still well below the minimum limit of 0.4 mm specified in BS EN 4678.

13

Porosity was observed at 2.4, 2.7 and 3.0 kW, all of which previously failed the BS EN 4678 criterion

on root width. Pores were found along the weld centreline with a diameter of 0.08, 0.14 and 0.08 mm

for 2.4, 2.7 and 3.0 kW, respectively, all smaller than the most stringent size of 0.40 mm specified in

BS EN 4678. Therefore, it was possible to obtain good quality welds at laser powers greater than 3.4

kW when welding at 6.0 m/min.

3.2 Effect of Welding Speed

Various welding speeds were also employed as listed in Table 5, now to examine the effect of welding

speed on the weld quality and microstructure. Fully penetrated welds were obtained at all welding

speeds from 1.8 to 4.2 m/min when using a fixed laser power of 2.0 kW. However, the root weld width

decreased with increasing welding speed. At fast welding speeds, a lack of penetration was observed

due to the low heat input arising from the fast welding speeds. This indicated that a careful balance

between the heat input and the welding speed was necessary. Figure 10 shows that the top and

bottom weld widths were inversely proportional to the welding speed. Both the top and bottom weld

widths were relatively wide at low welding speeds but the rate of change in the root width was greater.

The top width of all specimens satisfied the BS EN 4678 criterion whereas, the bottom width of the

specimens welded at 3.6 and 4.2 m/min did not. It was also found that changing the welding speed

influenced the weld pool geometry by affecting the liquid metal flow in the weld pool during welding,

and thus, produced different weld geometries as illustrated in Figure 11. A sharp decrease in the root

width caused the weld shape to change from an hourglass shape at low welding speeds to a nail

head shape above 3.0 m/min. The Rw as shown in Figure 10 decreases quickly with increasing

welding speed. This meant that the rate of change of the top width was once again relatively small

compared to that of the bottom width. Since the Rw dropped below 0.4 at above 3.6 m/min, the

welding process was considered unstable above this welding speed.

Undercut defects were detected both on the top and bottom surfaces but their depth was all less than

0.05 mm and therefore, acceptable according to all three criteria on the maximum tolerable depth of

undercut. The undercut depth initially increased with welding speed until 3.0 m/min, where it reached

the maximum and then decreased at higher welding speeds. Figure 10 suggests that smaller

undercuts were formed in nail head shaped welds due to the low heat input and less expulsion and

evaporation of molten materials at higher weld speeds. Similarly, the height of reinforcement

14

decreased with increasing welding speed, from around 0.20 mm at 1.8 m/min to 0.15 mm at 4.2

m/min. As mentioned previously, reinforcement was acceptable at all welding speeds, which would

otherwise increase non-value added costs. It was therefore, concluded that faster welding speeds are

preferred to minimise the levels of undercut as well as reinforcement, whereas, slower welding

speeds are preferred to obtain a stable full penetration

The level of porosity was dependent on the rate of weld pool solidification. At lower welding speeds,

the heat input increased and effectively reduced the cooling rate in the weld pool during solidification,

and thus, allowing gas bubbles to escape from the weld pool. As the weld pool remains in liquid state

for a longer period at lower welding speeds, it favours the nucleation, growth and finally the escape of

the gas bubbles. Keyhole instability and collapse may also contribute to the formation of porosity at

lower welding speeds (37). In contrast, at very high welding speeds, the solidification of the weld pool

is too fast that there is insufficient time for the growth of pores and therefore, the risk of porosity

formation is reduced. The chance of observing porosity is the greatest at intermediate speeds, where

there is enough time for pores to grow but also for the pool to solidify before the escape (45). In fact,

the largest pore was observed in the specimen welded at the welding speed of 3.0 m/min, with a

diameter of 0.23 mm, which was still less than the maximum pore size of 0.40 mm acceptable

according to BS EN 4678. As shown in Figure 11, porosity was observed over the entire range of

welding speeds used, all of them small enough to be acceptable and along the weld centreline at

lower half of the welds

Figure 12 shows the relationship between welding speed at fixed laser power of 2.5 kW, and the weld

widths. A very wide face and root widths were obtained at 1.8 m/min, with the top width exceeding the

limit of 3.2 mm in BS EN 4678 and the bottom width also being very close to the same upper limit of

3.2 mm. Compared to the specimens welded at 2.0 kW, the specimens welded at 2.5 kW showed a

steeper decrease in the top width with increasing welding speed, whereas, the bottom width

decreased with increasing welding speed at a similar rate as those welded at 2.0 kW. The face width

was too large at 1.8 m/min whereas, the root width was too narrow at 4.2 and 4.8 m/min. The

transverse section of the specimen welded at 1.8 m/min in Figure 13 has a noticeably wide face and

root weld widths with almost rectangular weld shape. The specimen welded at a faster welding speed

of 2.4 m/min displayed an hourglass shape, which changed into a nail head shape at 4.2 m/min where

the root width became too narrow according to the 1.0 mm minimum root width specified in BS EN

15

4678. The trend in Rw was the same as that for the specimens welded at 2.0 kW, where it dropped

below the limit of 0.4 at the same welding speed of 4.2 m/min. The specimen welded at 1.8 m/min

produced undercut with a depth greater than the maximum undercut depth of 0.10 mm cited in the

less stringent BS EN ISO 13919-2 and BS EN 4678. The specimens welded at 2.4 and 3.0 m/min

pass the criterion of BS EN ISO 13919-2 and BS EN 4678 but failed that of AWS D17.1. The

specimens welded at higher welding speeds of 3.6, 4.2 and 4.8 m/min satisfied all criteria with

undercut depths less than 0.05 mm. The height of reinforcement in specimens were all below 0.4 mm,

with a maximum of around 0.25 mm, so they were all acceptable. The reinforcement decreased with

increasing welding speed. Increasing the welding speed had a positive influence on reducing the

extent of imperfections in the welds, whereas, sufficiently fast welding speed was necessary to

achieve a stable full penetration.

Porosity was observed in every specimen for the welding speeds investigated at the given constant

power of 2.5 kW. The pores were spherical in shape and their diameters were 0.04, 0.08, 0.04, 0.05,

0.14 and 0.10 mm for the welding speed of 1.8, 2.4, 3.0, 3.6, 4.2 and 4.8 m/min, respectively. The

pores were the largest at the two fastest welding speeds of 4.2 and 4.8 m/min with nail head shaped

welds. The narrow root width at higher speeds increased the probability of retaining porosity

(43,45,49), because the escape of gas porosity before solidification became more difficult and only

accessible via the top surface (48). The low Rw at 4.2 and 4.8 m/min implied that keyhole instability

and collapse at very high welding speeds contributed to increased pore diameters. Nevertheless, the

size of all pores were confidently within the safe limit of 0.40 mm specified in BS EN 4678.

According to Figure 13, when a fixed laser power of 3.0 kW was used, the operating window for

welding speed became wider as shown by the weld widths of the specimens welded at welding

speeds of 4.2, 4.8 and 5.4 m/mins, those of which at previously investigated lower laser powers failed

to meet the minimum root width of 1.0 mm specified in BS EN 4678. In this case, the top weld width

was acceptable over the entire welding speed range used. The bottom weld width of all specimens

except at 6.0 m/min, were acceptable. For this set of welding speeds and laser power, the Rw values

were all above 0.4. In other words, a stable full penetration was obtained under all conditions. The

weld cross-section exhibited an hourglass shape at low welding speeds up to around 3.6 m/min and

then changed to a nail head shape with some features resembling the hourglass shape. Undercut

was not an issue except at 2.4 m/min which was very close to the limiting depth. None of the

16

specimens passed the most stringent depth of 0.05 mm specified in AWS D17.1. The influence of

welding speed on reinforcement was obvious, with a steady reduction in the height of the

reinforcement with increasing welding speed. Overall, the amount of reinforcement in these

specimens were well acceptable.

Porosity was not observed at all welding speeds from 1.8 to 6.0 m/min. Except the limiting maximum

welding speed of 6.0 m/min, which produced too narrow root width, any welding speed below 6.0

m/min can therefore, be used to produce a good quality weld at the given fixed laser power of 3.0 kW,

which also meant that a wide operating window was obtained.

A further reduction in the overall weld widths with increasing welding speed was observed as shown

in Figure 16 at the given constant laser power of 3.5 kW, than those of 2.5 and 3.0 kW. The widest

top and bottom weld widths were produced at 2.4 m/min. The top and the bottom weld widths only

varied by around 0.2 mm from 3.6 m/min to 7.2 m/min, with values close to 2.0 mm and 1.0 mm,

respectively. The bottom widths at welding speeds greater than 2.4 m/min were very close to the

minimum limit of 1.0 mm cited in BS EN 4678. The top widths, on the other hand, were at least 1 mm

less than the maximum limit of 3.2 mm. The weld shape at 2.4 m/min was an hourglass shape but nail

head shape above 3.6 m/min, with considerably narrow root widths. The Rw shown in Figure 16 was

the largest at 2.4 m/min, as proved by its hourglass weld shape, and then decreased to around 0.6 for

the specimens welded at higher welding speeds as shown by their nail head shaped welds. With the

Rw greater than 0.4, the welding process produced a stable full penetration for all welding speeds

investigated. For all the specimens welded at 3.5 kW, the undercut remained fairly constant

regardless of welding speed, with values within both the less stringent 0.10 mm and the more

stringent 0.05 mm limits. Reinforcement or excessive root penetration was observed at 2.4 m/min with

a height of around 0.25 mm at bottom. Reinforcement decreased with increasing welding speed and

its effect on weld quality was insignificant and well below 0.4 mm limit. Porosity was only observed at

the fastest welding speed of 7.2 m/min. However, the size of 0.04 mm was almost an order of

magnitude smaller than the maximum tolerable size of 0.4 mm specified in BS EN 4678. Interestingly,

for this particular specimen, a spatter was found at the bottom surface near the fusion boundary but

there were no welding quality acceptance criteria on spatter according to AWS D17.1, BS EN ISO

13919-2 and BS EN 4678, except in BS EN ISO 13919-2 which states that the acceptance of spatter

17

depends on applications. The effect of spatter on weld quality or performance was therefore,

considered negligible.

3.3 Effect of Focal Position

Finally, various focal positions were employed as listed in Table 6 to examine the effect of changing

focal position on the weld quality and microstructure. Figure 18 shows the effect of varying focal

distance on the weld width, which is the distance relative to the surface of the workpiece, with a value

equal to zero on the surface, positive when the beam is focused above the surface of the workpiece

and negative below. The location of the focal position determines the intensity of the beam on the

surface of the workpiece. The laser beam energy density was the maximum at the 0 mm focal point.

By defocusing the beam either above or below the top surface of the workpiece, it was possible to

change the beam intensity and produce a wider weld pool size and larger keyhole which were

sometimes beneficial for processing stability. On the other hand, it was also desired to have a small

focal spot as it meant that there is less material involved during welding. Positive defocusing reduced

the irradiance on the surface but the keyhole was inherently more stable. Unlike Pastor et al. (48) who

observed that positive defocusing leads to more plasma which defocuses the beam, this was not a

serious problem with fibre laser as the plasma effect was found to be very small with fibre laser due to

its wavelength. Negative defocusing led to an increased laser beam coupling to the material and

increased the irradiance inside the weld pool (51) so the welding speed could be increased compared

to without defocusing (52). It was therefore, important to keep the focal distance within the appropriate

range to maintain the required power density to form a keyhole. The optimum focal position resulted

in the maximum Rw, suitable face and root widths and minimum imperfections.

Full penetration was achieved in all welds regardless of the focal position used as shown in Figure 18.

The top and the bottom weld widths were measured to be the minimum at 0 mm as expected. The

general trend was that the weld widths increased with the magnitude of focal position. As the

threshold power density for keyhole formation was lower at negative focal positions due to improved

laser and material coupling within the material (50), the weld widths were found to be larger at

negative focal positions than at positive focal positions of the same magnitude. The BS EN 4678

criteria on face and root widths, and the Rw limit of 0.4 were satisfied at all focal positions. The Rw was

calculated to be greater at ±2 mm and 0 mm than at ±4 and ±5.5 mm. Interestingly, for smaller focal

18

positions, the weld cross-sections displayed an hourglass shape, whereas, at larger focal positions, it

was a V shape. Undercut defect was found to be acceptable at all focal positions due to the fact that

the chosen laser power and the welding speed values have already been optimised as discussed

above and also partly because the heat input was not affected by the focal position, while only the

laser power density changed. The depth of undercut in all specimens satisfied the BS EN 4678 and

BS EN ISO 13919-2 criteria. The specimens welded at +2 and -4 mm focal positions showed the

largest undercut but were still acceptable according to these criteria. The smallest undercut was

produced at +5.5 mm which satisfied the AWS D17.1 criterion. The size of undercut measured both in

the specimens welded at 0 and +4 mm was also small. There was only a small change in the height of

reinforcement with focal position except the specimen welded at a focal position of 5.5 mm which

exhibited excessive penetration due to the effect of laser-material interaction discussed above when

negatively defocused.

Figure 19 shows that porosity was formed at focal positions of -4.0 and +5.5 mm and the size of pores

were very small at 0.04 and 0.03 mm respectively. Porosity was found in the specimens with small Rw

values, where a small Rw is equivalent to a large aspect ratio (53). Therefore, it was determined that

overall the most acceptable weld profile was obtained at a focal position of +4 mm because a stable

full penetration was achieved with a sufficiently wide root width, small under and reinforcement and

finally, the absence of porosity defects. The specimen welded at -2 mm would also produce a good

quality weld for the same reason discussed above. Note that the presence of porosity defect was

observed in the lower half of the FZ. The spherical shape of the pores indicate that they were caused

by entrapment of hydrogen gas bubble in the molten weld pool during solidification due to increased

solubility of hydrogen with increasing temperature in Ti-6Al-4V.

3.4 Relationship between weld microstructure and process parameters

The weld microstructure was found to be directly related to heat input. The heat input was controlled

by changing laser power and/or welding speed, where higher heat input was achieved by either

increasing the laser power or decreasing the welding speed, and vice versa. As Figure 20 shows, for

lower heat inputs, as achieved by increasing the welding speed or decreasing the laser power, the

time spent at higher temperatures decreased and the cooling rate increased over the solidification

temperature range, arising from a smaller weld pool volume. This promoted the formation of finer

martensite and prior β grains in the FZ microstructure. On the other hand, for higher heat inputs, as

19

achieved by decreasing the welding speed or increasing the laser power, the time spent at high

temperatures increased and the cooling rate decreased. As a result, the martensitic phase changed

from a needle like to plate like morphology and the prior β grain size increased due to longer time

available for grain coarsening. In addition, diffusional transformation constituents such as acicular α

and grain boundary α appeared.

Changing the focal position also changed the incident laser power density but not the heat input.

When the focal position was set to 0 mm, higher power density or longer interaction time was

obtained, which promoted additional grain growth in the fusion zone. However, since the change in

focal position was relatively small, no significant differences in microstructure was observed. Instead,

it mainly affected the weld width, where increasing or decreasing the focal position also increased

both the face and the root weld widths.

4. Conclusions

Thin sheets of titanium alloy Ti-6Al-4V were fibre laser welded to investigate the influence of laser power,

welding speed and beam focal position on the weld quality, shape, microstructure and defects. The

following conclusions can be drawn.

Rapid heating and cooling of fibre laser welding induced a fine martensitic microstructure in the FZ and

mostly martensitic and acicular α structure in the adjacent HAZ. The HAZ next to the BM consisted of

primary α, intergranular β and martensite. The BM consisted of primary α and intergranular β.

Heat input supplied to the workpiece was mainly controlled by laser power and welding speed, which

in turn, influenced the weld microstructure by modifying the peak temperatures experienced and, the

heating and cooling rates. Decreasing the laser power or increasing the welding speed resulted in finer

martensite and prior β grains in the FZ, whereas, increasing the laser power or decreasing the welding

speed resulted in acicular martensitic structure and larger prior β grain size, as well as the formation of

diffusional transformation constituents such as acicular α and grain boundary α. No significant

differences in weld microstructure were observed over the range of focal positions investigated.

The weld top and bottom width all increase with increasing laser power, focal position and decreasing

welding speed.

20

Incomplete penetration or narrow root width were the main problems at low laser powers and fast

welding speeds, whereas, undercut was the main defect at high laser powers.

In general, the weld quality produced was good for all welding speeds as long as it was not too fast.

Spatter was observed at the bottom surface at very fast welding speeds.

Reinforcement increased with increasing laser power or welding speed but was not critical enough to

affect the weld quality. Porosity was detected were all below the critical size

The root width increased more at negative focal positions than at the equivalent positive focal positions

due to an enhanced interaction between laser and the material being welded with negative defocusing.

It was therefore, possible to weld Ti-6Al-4V at high laser powers and welding speeds, over a wide range

of focal positions, with characteristics and qualities that are pertinent to aerospace applications.

5. Acknowledgement

The strong support from the Aviation Industry Corporation of China (AVIC) and Beijing Aeronautical

Manufacturing Technology Research Institute (BAMTRI) for this funded research is much

appreciated. The research was performed at the AVIC Centre for Structural Design and Manufacture

at Imperial College London. Dr C M Davies acknowledges the support of EPSRC under grant number

EP/I004351/1.

21

6. References

1. Clifford M. ASME engineer’s data book. 1st ed. Fairfield: ASME Press; 2001. 211 p.

2. Cao X, Wallace W, Poon C, Immarigeon JP. Research and Progress in Laser Welding of Wrought Aluminum Alloys. I. Laser Welding Processes. Materials and Manufacturing Processes. 2003;18(1): 1–22.

3. Benyounis KY, Olabi AG, Hashmi MSJ. Effect of laser welding parameters on the heat input and weld-bead profile. Journal of Materials Processing Technology. 2005;164-165: 978–985.

4. El-Batahgy A-M, Kutsuna M. Laser Beam Welding of AA5052, AA5083, and AA6061 Aluminum Alloys. Research Letters in Materials Science. 2009;2009: 1–9.

5. El-Batahgy A-M. Effect of laser welding parameters on fusion zone shape and solidification structure of austenitic stainless steels. Materials Letters. 1997;32(2-3): 155–163.

6. Gong S, Chen L, Yao W, Tang T. Research on laser welding of BT20 titanium alloy. In: Li L, Achara C (eds.) Proceedings of the 22nd International Congress on Application of Lasers & Electro-Optics. Jacksonville, Fl: Laser Institute of America; 2003. p. 1–6.

7. Chen L, Shuili G, Yang J. Study on Full Penetration Stability of Light Alloys Sheet Laser Welding. In: Hinduja S, Li L (eds.) Proceedings of the 37th International MATADOR Conference. Manchester: University of Manchester; 2012. p. 315–318.

8. Chen L, Hu L, Gong S. A study on the porosity of CO2 laser welding of titanium alloy. China Welding. 2006;15(1): 1–5.

9. Karlsson J, Kaplan AFH. Analysis of a fibre laser welding case study, utilising a matrix flow chart. Applied Surface Science. 2011;257(9): 4113–4122.

10. Karlsson J, Kaplan A. Fibre laser welding for lightweight design. ICALEO 2009 - International Congress on Applications of Laser and Electro-Optics. 2009. p. 1548–1557.

11. Karlsson J, Ilar T, Kaplan AFH. Knowledge platform approach for fiber laser welding of high strength. Proceedings of the 12th NOLAMP conference. Copenhagen: FORCE Technology; 2009. p. 1–15.

12. Norman PM, Kaplan AFH, Karlsson J. Classification and generalization of data from a fibre-laser hybrid welding case. Physics Procedia. 2010;5: 69–76.

13. Karlsson J, Markmann C, Minhaj Alam M, Kaplan AFH. Parameter influence on the laser weld geometry documented by the matrix flow chart. Physics Procedia. 2010;5: 183–192.

14. Manonmani K, Murugan N, Buvanasekaran G. Effect of process parameters on the weld bead geometry of laser beam welded stainless steel sheets. International Journal for the Joining of Materials. 2005;17(4): 103–109.

15. Balasubramanian TS, Balakrishnan M, Balasubramanian V, Muthu Manickam MA. Influence of welding processes on microstructure, tensile and impact properties of Ti-6Al-4V alloy joints. Transactions of Nonferrous Metals Society of China (English Edition). The Nonferrous Metals Society of China, 2011;21(6): 1253–1262.

16. Squillace A, Prisco U, Ciliberto S, Astarita A. Effect of welding parameters on morphology and mechanical properties of Ti-6Al-4V laser beam welded butt joints. Journal of Materials Processing Technology. 2012;212(2): 427–436.

17. Campanelli SL, Casalino G, Mortello M, Angelastro A, Ludovico AD. Microstructural Characteristics and Mechanical Properties of Ti6Al4V Alloy Fiber Laser Welds. Procedia CIRP. 2015;33: 428–433.

22

18. Mueller S, Stiles E, Dienemann R. Study of porosity formation during laser welding of Ti6Al4V. Proceedings of the 25th International Congress on Applications of Lasers and Electro-Optics. Scottsdale, AZ: Laser Institute of America; 2006. p. 304.

19. Mueller S, Braat C, Mueller P, Cuddy J, Shankar K. Laser beam welding of titanium – a comparison of CO2 and fibre laser for potential aerospace applications. Proceedings of the 27th International Congress on Applications of Lasers and Electro-Optics. Temecula: Laser Institute of America; 2008. p. 1802.

20. Hilton P, Blackburn J, Chong P. Welding of Ti-6Al-4V with fibre delivered laser beams. Proceedings of the 26th International Congress on Applications of Lasers and Electro-optics. Orlando, FL: Laser Institute of America; 2007. p. 887–895.

21. Tsukamoto S, Steen WM, Mazumder J. Laser Welding. Laser Material Processing. 4th ed. London: Springer-Verlag London; 2010. p. 199–249.

22. BS EN 4678:2011 Aerospace series. Weldments and brazements for aerospace structures. Joints of metallic materials by laser beam welding. Quality of weldments.

23. ISO/TR17671-6:2005 Welding - Recommendations for welding of metallic materials - Part 6: Laser beam welding.

24. AWS C7.2 Recommended Practices for Laser Beam Welding, Cutting and Drilling.

25. BS EN 1011-6:2005 Welding. Recommendations for welding of metallic materials. Laser beam welding.

26. ASTM E3 Methods for Preparation of Metallographic Specimens.

27. BS EN 1321:1997 Destructive test on welds in metallic materials. Macroscopic and microscopic examination of welds.

28. Peters M, Hemptenmacher J, Kumpfert J, Leyens C. Structure and Properties of Titanium and Titanium Alloys. In: Leyens C, Peters M (eds.) Titanium and Titanium Alloys. Fundamentals and Applications. 1st ed. Weinheim: WILEY-VCH Verlag GmbH & Co. KGaA; 2003. p. 1–36.

29. BS EN ISO 15614-11:2002 Specification and qualification of welding procedures for metallic materials — Welding procedure test — Part 11 Electron and laser beam welding.

30. ISO 13919-2 Welding — Electron and laser-beam welded joints — Guidance on quality levels for imperfections — Part 2: Aluminium and its weldable alloys.

31. AWS D17.1:2001 Specification for Fusion Welding for Aerospace Applications.

32. Mendez PF, Eagar TW. Welding processes for aeronautics. Advanced Materials & Processes. 2001;159(5): 39–43.

33. Elmer JW, Palmer TA, Babu SS, Zhang W, DebRoy T. Phase transformation dynamics during welding of Ti-6Al-4V. Journal of Applied Physics. 2004;95: 8327–8339.

34. Rai R, Elmer JW, Palmer T a, DebRoy T. Heat transfer and fluid flow during keyhole mode laser welding of tantalum, Ti–6Al–4V, 304L stainless steel and vanadium. Journal of Physics D: Applied Physics. 2007;40(18): 5753–5766.

35. Wang SHH, Wei MDD, Tsay LWW. Tensile properties of LBW welds in Ti–6Al–4V alloy at evaluated temperatures below 450 °C. Materials Letters. 2003;57(12): 1815–1823.

36. Elmer JW, Palmer TA, Babu SS, Zhang W, DebRoy T. Phase transformation dynamics during welding of Ti-6Al-4V. Journal of Applied Physics. 2004;95: 8327–8339.

37. Kabir ASH, Cao X, Medraj M, Wanjara P, Cuddy J, Birur A. Effect of Welding Speed and Defocusing Distance on the Quality of Laser Welded Ti-6Al-4V. Proceedings of the Materials Science and Technology (MS&T) 2010 Conference. Houston, TX: ASM International; 2010. p. 2787–2797.

23

38. Kabir ASH, Cao X, Gholipour J, Wanjara P, Cuddy J, Birur A, et al. Effect of Postweld Heat Treatment on Microstructure, Hardness, and Tensile Properties of Laser-Welded Ti-6Al-4V. Metallurgical and Materials Transactions A. 2012;43(11): 4171–4184.

39. Ahmed T, Rack HJ. Phase transformations during cooling in α+β titanium alloys. Materials Science and Engineering: A. 1998;243: 206–211.

40. Babu B. Physically Based Model for Plasticity and Creep of Ti-6Al-4V. Luleå University of Technology; 2008.

41. Cao X, Jahazi M. Effect of welding speed on butt joint quality of Ti–6Al–4V alloy welded using a high-power Nd:YAG laser. Optics and Lasers in Engineering. 2009;47(11): 1231–1241.

42. Kabir ASH. Nd:YAG Laser Welding of Ti-6Al-4V Alloy. Concordia University; 2011.

43. Murav’ev VI. Problems of Pore Formation in Welded Joints of Titanium Alloys. Metal Science and Heat Treatment. 2005;47(7-8): 282–288.

44. Fan Y, Chen Z, Zhang CH, Liu AM. A comparison of microstructure and mechanical properties of welded thin Ti6Al4V with three different types of laser. Materials Research Innovations. 2015;19(S4): S187–S192.

45. Khaled T. An investigation of pore cracking in titanium welds. Journal of Materials Engineering and Performance. 1994;3(1): 21–36.

46. Mitchell DR. Porosity in Titanium Welds. Welding Journal. 1982;61: 157s – 167s.

47. Verhaeghe G, Hilton P, Barnes S. Achieving Low-Porosity Laser Welds in Aerospace Aluminium Alloy. 2003 Aerospace Manufacturing Technology Conference. Montreal: SAE International; 2003. p. 2003–01 – 2895.

48. Pastor M, Zhao H, Debroy T. Pore formation during continuous wave Nd:YAG laser welding of aluminium for automotive applications. Welding International. p, 2001;15(4): 275–281.

50. Zhou J, Tsai H-L. Porosity Formation and Prevention in Pulsed Laser Welding. Journal of Heat Transfer. 2007;129(8): 1014.

51. Zhao H, Martukanitz RP, Debroy T. Porosity, Underfill and Magnesium Loss during Continuous Wave Nd:YAG Laser Welding of Thin Plates of Aluminum Alloys 5182 and 5754. Welding Journal. 1999;78(6): 207–216.

52. Leong KH, Sabo KR, Sanders PG, Spawr WJ. Laser Welding of Aluminum Alloys. Proceedings of Photonics West ’97. International Society for Optics and Photonics; 1997. p. 37–44.

53. Schuöcker D. Welding with Lasers. High Power Lasers in Production Engineering. London: Imperial College Press; 1999. p. 337–370.

54. Kim JS, Watanabe T, Yoshida Y. Effect of the beam-defocusing characteristics on porosity formation in laser welding. Journal of Materials Science Letters. 1995;14(22): 1624–1626.

24

Table 1 Chemical composition of Ti-6Al-4V Grade 5 (Wt. %)

Ti Al V Fe O

Balance 5.5-6.76 3.5-4.5 0.25 0.2

Table 2 Examination and tests for welds in accordance with acceptance level D specified in BS EN ISO 15614-11:2002 (29)

Test piece Type of examination and test Extent of examination and test

Butt weld

Visual examination Radiographic examination Ultrasonic examination Surface crack detection Metallographic examination

100 % if required if required if required 1 section minimum

T-joint

Visual examination Ultrasonic examination Surface crack detection Metallographic examination

100 % if required if required 1 section minimum

Table 3 Relevant welding quality acceptance criteria from AWS D17.1:2001, BS EN ISO 13919-2-2001 and BS EN 4678:2011 standards (22,30-31)

Standard Level Face width (mm)

Root width (mm)

Porosity (mm)

Undercut (mm)

Reinforcement (mm)

AWS D17.1 Class A N/A N/A ≤ 0.66 ≤ 0.05 ≤ 0.66 BS EN ISO 13919-2 stringent B N/A N/A ≤ 0.60 ≤ 0.10 ≤ 0.50 BS EN 4678 Titanium alloys ≤ 3.20 ≤ 1.00 ≤ 0.40 ≤ 0.10 ≤ 0.40

25

Table 4 Process parameters tested to investigate the effect of laser power and the resulting

measured top and bottom weld widths and calculated Rw

Laser power (kW)

Welding speed

(m/min)

Focal position

(mm)

Top width (mm)

Bottom width (mm)

Rw

3.6 2.1 4.0 2.64 1.82 0.69 3.3 2.1 4.0 2.77 2.04 0.74 3.0 2.1 4.0 2.77 2.08 0.75 2.7 2.1 4.0 2.78 2.12 0.76 2.4 2.1 4.0 2.94 2.30 0.78 2.1 2.1 4.0 2.80 2.35 0.84 1.8 2.1 4.0 2.49 2.03 0.82 1.5 2.1 4.0 2.18 1.45 0.67 1.2 2.1 4.0 1.99 0.00 0.00 4.5 3.9 4.0 2.09 1.37 0.66 4.2 3.9 4.0 2.09 1.33 0.64 3.9 3.9 4.0 2.15 1.37 0.64 3.6 3.9 4.0 2.14 1.28 0.60 3.3 3.9 4.0 2.12 1.35 0.64 3.0 3.9 4.0 2.00 1.17 0.59 2.7 3.9 4.0 1.92 0.91 0.47 2.4 3.9 4.0 1.79 0.59 0.33 2.1 3.9 4.0 1.60 0.00 0.00 1.8 3.9 4.0 1.50 0.00 0.00 4.5 6.0 4.0 1.81 1.12 0.62 4.2 6.0 4.0 1.80 1.11 0.62 3.9 6.0 4.0 1.77 1.12 0.63 3.6 6.0 4.0 1.82 1.10 0.60 3.4 6.0 4.0 1.91 1.09 0.57 3.0 6.0 4.0 1.79 0.86 0.48 2.7 6.0 4.0 1.64 0.60 0.37 2.4 6.0 4.0 1.69 0.06 0.04 2.1 6.0 4.0 1.50 0.00 0.00

26

Table 5 Process parameters tested to investigate the effect of welding speed and the resulting

measured top and bottom weld widths and calculated Rw

Laser power (kW)

Welding speed

(m/min)

Focal position

(mm)

Top width (mm)

Bottom width (mm)

Rw

2.0 1.8 4.0 2.77 2.51 0.91 2.0 2.4 4.0 2.07 1.14 0.55 2.0 3.0 4.0 2.28 1.76 0.77 2.0 3.6 4.0 1.84 0.70 0.38 2.0 4.2 4.0 1.76 0.21 0.12 2.5 1.8 4.0 3.38 2.90 0.85 2.5 2.4 4.0 2.66 2.19 0.82 2.5 3.0 4.0 2.24 1.60 0.71 2.5 3.6 4.0 2.03 1.10 0.54 2.5 4.2 4.0 1.89 0.70 0.37 2.5 4.8 4.0 1.78 0.40 0.22 3.0 1.8 4.0 2.53 1.92 0.76 3.0 2.4 4.0 2.41 1.89 0.78 3.0 3.0 4.0 1.96 1.30 0.66 3.0 3.6 4.0 2.05 1.32 0.64 3.0 4.2 4.0 1.89 1.08 0.57 3.0 4.8 4.0 1.95 1.11 0.57 3.0 5.4 4.0 1.73 1.10 0.64 3.0 6.0 4.0 1.73 0.87 0.50 3.5 2.4 4.0 2.26 1.86 0.82 3.5 3.6 4.0 1.96 1.20 0.61 3.5 4.8 4.0 1.78 1.10 0.62 3.5 6.0 4.0 1.77 0.99 0.56 3.5 7.2 4.0 1.82 1.00 0.55

Table 6 Process parameters tested to investigate the effect of focal position and the resulting

measured top and bottom weld widths and calculated Rw

Laser power (kW)

Welding speed

(m/min)

Focal position

(mm)

Top width (mm)

Bottom width (mm)

Rw

2.5 3.0 0.0 1.53 1.10 0.72 2.5 3.0 4.0 2.20 1.22 0.55 2.5 3.0 2.0 1.85 1.46 0.79 2.5 3.0 5.5 2.24 1.23 0.55 2.5 3.0 -2.0 1.92 1.44 0.75 2.5 3.0 -4.0 2.59 1.69 0.65 2.5 3.0 -5.5 2.24 1.44 0.64

27

List of tables

Table 1 Chemical composition of Ti-6Al-4V Grade 5 (Wt. %)

Table 2 Examination and tests for welds in accordance with acceptance level D specified in BS EN ISO

15614-11:2002 (3)

Table 3 Relevant welding quality acceptance criteria from AWS D17.1:2001, BS EN ISO 13919-2-2001

and BS EN 4678:2011 standards (13-15)

Table 4 Process parameters tested to investigate the effect of laser power and the resulting measured

top and bottom weld widths and calculated Rw

Table 5 Process parameters tested to investigate the effect of welding speed and the resulting

measured top and bottom weld widths and calculated Rw

Table 6 Process parameters tested to investigate the effect of focal position and the resulting measured

top and bottom weld widths and calculated Rw

List of figures

Figure 1 Experimental set-up for fibre laser welding Ti-6Al-4V sheets

Figure 2 Microstructure of fibre laser welded Ti-6Al-4V (P=2.7 kW, V=2.1 m/min and f=+4mm) in a) the base metal, b) the heat affected zone and c) the fusion zone at 200x and 500x magnifications etched using Kroll’s reagent

Figure 3 a) Energy-dispersive X-ray spectroscopy (EDS) spectrum of the fusion zone and b) the detected weight percentage (%) of chemical elements in the parent material, the heat affected zone and the fusion zone

Figure 4 a) Relationship between weld width and laser power at a welding speed of 2.1 m/min with +4 mm defocus and argon shielding gas and b) the resultant weld width ratio, undercut and reinforcement

Figure 5 Transverse sections of welds produced with different laser powers at a welding speed of 2.1 m/min with +4 mm defocus and argon shielding gas

Figure 6 a) Relationship between weld width and laser power at a welding speed of 3.9 m/min with +4 mm defocus and argon shielding gas and b) the resultant weld width ratio, undercut and reinforcement

Figure 7 Transverse sections of welds produced with different laser powers at a welding speed of 3.9 m/min with +4 mm defocus and argon shielding gas

Figure 8 a) Relationship between weld width and laser power at a welding speed of 6.0 m/min with +4 mm defocus and argon shielding gas and b) the resultant weld width ratio, undercut and reinforcement

Figure 9 Transverse sections of welds produced with different laser powers at a welding speed of 6.0 m/min with +4 mm defocus and argon shielding gas

Figure 10 a) Relationship between weld width and welding speed at a laser power of 2.0 kW with +4 mm defocus and argon shielding gas and b) the resultant weld width ratio, undercut and reinforcement

Figure 11 Transverse sections of welds produced with different welding speeds at a laser power of 2.0 kW with +4 mm defocus and argon shielding gas

Figure 12 a) Relationship between weld width and welding speed at a laser power of 2.5 kW with +4 mm defocus and argon shielding gas and b) the resultant weld width ratio, undercut and reinforcement4.2

28

Figure 13 Transverse sections of welds produced with different welding speeds at a laser power of 2.5 kW with +4 mm defocus and argon shielding gas

Figure 14 a) Relationship between weld width and welding speed at a laser power of 3.0 kW with +4 mm defocus and argon shielding gas and b) the resultant weld width ratio, undercut and reinforcement

Figure 15 Transverse sections of welds produced with different welding speeds at a laser power of 3.0 kW with +4 mm defocus and argon shielding gas

Figure 16 a) Relationship between weld width and welding speed at a laser power of 3.5 kW with +4 mm defocus and argon shielding gas and b) the resultant weld width ratio, undercut and reinforcement

Figure 17 Transverse sections of welds produced with different welding speeds at a laser power of 3.5 kW with +4 mm defocus and argon shielding gas

Figure 18 a) Relationship between weld width and focal position at a laser power of 2.5 kW, a welding speed of 3.0 m/min and with argon shielding gas, and b) the resultant weld width ratio, undercut and reinforcement

Figure 19 Transverse sections of welds produced with different focal positions at a laser power of 2.5 kW, a welding speed of 3.0 m/min and with argon shielding gas

Figure 20 Comparison of microstructures in the fusion zone obtained using four different combinations of welding speed and laser power

KUKA

Robert Fiber

Laser

Laser welding head

Gas

knife

Protecting

Gas

Shielding

Gas

Welding fixture

Welding

sheet

Control system

FZ

BM

HAZ

P=2.7 kW, V=2.1 m/min, f=+4 mm

10 mm 100 μm

100 μm 100 μm

a) b)

89.94

5.43 4.63

89.78

5.30 4.92

89.43

5.12 5.45

0

10

20

30

40

50

60

70

80

90

100

Ti Al V

We

ight (%

) PM

HAZ

FZ

a) b)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0 1 2 3 4 5

Weld

wid

th (

mm

)

Laser power (kW)

Top

Bot

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 1 2 3 4 5

Imp

erf

ectio

ns (

mm

)

Rw

Laser power (kW)

Width ratio

Undercut

Reinforcement

1.2 kW 1.5 kW 1.8 kW 2.1 kW 2.4 kW

2.7 kW 3.0 kW 3.3 kW 3.6 kW

1.0 mm 1.0 mm 1.0 mm 1.0 mm

1.0 mm 1.0 mm 1.0 mm 1.0 mm 1.0 mm

a) b)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0 1 2 3 4 5

We

ld w

idth

(m

m)

Laser power (kW)

Top

Bot0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 1 2 3 4 5

Imperf

ections (

mm

)

Rw

Laser power (kW)

Width ratio

Undercut

Reinforcement

1.8 kW 2.1 kW 2.4 kW 2.7 kW 3.0 kW

3.3 kW 3.6 kW 3.9 kW 4.2 kW 4.5 kW

1.0 mm 1.0 mm 1.0 mm

1.0 mm 1.0 mm 1.0 mm

1.0 mm

1.0 mm

1.0 mm

1.0 mm

a) b)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0 1 2 3 4 5

We

ld w

idth

(m

m)

Laser power (kW)

Top

Bot

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 1 2 3 4 5

Imp

erf

ectio

ns (

mm

)

Rw

Laser power (kW)

Width ratio

Undercut

Reinforcement

2.1 kW 2.4 kW 2.7 kW 3.0 kW 3.4 kW

3.6 kW 3.9 kW 4.2 kW 4.5 kW

1.0 mm 1.0 mm 1.0 mm 1.0 mm 1.0 mm

1.0 mm 1.0 mm 1.0 mm 1.0 mm

a) b)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0 1 2 3 4 5 6 7 8

We

ld w

idth

(m

m)

Speed (m/min)

Top

Bot0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 1 2 3 4 5 6 7 8

Imp

erf

ectio

ns (

mm

)

Rw

Speed (m/min)

Width ratio

Undercut

Reinforcement

1.8 m/min 2.4 m/min 3.0 m/min 3.6 m/min 4.2 m/min

1.0 mm 1.0 mm 1.0 mm 1.0 mm 1.0 mm

a) b)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0 1 2 3 4 5 6 7 8

We

ld w

idth

(m

m)

Speed (m/min)

Top

Bot0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 1 2 3 4 5 6 7 8

Imp

erf

ectio

ns (

mm

)

Rw

Speed (m/min)

Width ratio

Undercut

Reinforcement

1.8 m/min 2.4 m/min 3.0 m/min

3.6 m/min 4.2 m/min 4.8 m/min

1.0 mm 1.0 mm

1.0 mm

1.0 mm

1.0 mm 1.0 mm

a) b)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0 1 2 3 4 5 6 7 8

We

ld w

idth

(m

m)

Speed (m/min)

Top

Bot0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 1 2 3 4 5 6 7 8

Imp

erf

ectio

ns (

mm

)

Rw

Speed (m/min)

Width ratio

Undercut

Reinforcement

1.8 m/min 2.4 m/min 3.0 m/min 3.6 m/min

4.2 m/min 4.8 m/min 5.4 m/min 6.0 m/min

1.0 mm

1.0 mm

1.0 mm

1.0 mm

1.0 mm

1.0 mm

1.0 mm

1.0 mm

a) b)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0 1 2 3 4 5 6 7 8

We

ld w

idth

(m

m)

Speed (m/min)

Top

Bot

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 1 2 3 4 5 6 7 8

Imperf

ections (

mm

)

Rw

Speed (m/min)

Width ratio

Undercut

Reinforcement

2.4 m/min 3.6 m/min 4.8 m/min 6 m/min 7.2 m/min

1.0 mm 1.0 mm 1.0 mm 1.0 mm 1.0 mm

a) b)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

-6 -4 -2 0 2 4 6

We

ld w

idth

(m

m)

Focal position (mm)

Top

Bot

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.0

0.2

0.4

0.6

0.8

1.0

1.2

-6 -4 -2 0 2 4 6

Imperf

ections (

mm

)

Rw

Focal position (mm)

Width ratio

Undercut

Reinforcement

+2.0 mm +4.0 mm +5.5 mm

+0.0 mm

-5.5 mm -4.0 -2.0 mm

1.0 mm

1.0 mm 1.0 mm 1.0 mm

1.0 mm 1.0 mm 1.0 mm

mm

100 μm 100 μm

100 μm100 μm

P=2.1 kW, V=2.1 m/min, f=+4 mm P=4.2 kW, V=2.1 m/min, f=+4 mm

P=2.1 kW, V=6.0 m/min, f=+4 mm P=4.2 kW, V=6.0 m/min, f=+4 mm

Change in laser powerC

han

ge in

wel

din

g sp

eed

100 μm 100 μm

100 μm 100 μm