To analyze the weld geometry of the sample, the average weld width and penetration depth were measured. The effects of the center beam location on the bond width and penetration depth are shown in Figs.5.3 and 5.4. The weld bond width and penetration depth decreased significantly with increasing distance of the beam center location from the edge of the upper specimen. The weld bond width and penetration depth decreased significantly with increasing distance of the beam center from the edge of the upper specimen. A larger distance results in a low heat concentration on the upper surface of the specimen. This occurs due to heat being dissipated to the surroundings because the magnesium alloy exhibits high thermal conductivity. Less heat is absorbed by the workpiece, resulting in a smaller melted weld area. An increased distance may also increase the size of voids at the root, as shown in Fig. 5.2. From this result, we conclude that the weld geometry in laser welding is very sensitive to beam center location distances on the order of 100 µm.

#### **5.2.3 Microstructure**

148 Nd YAG Laser

width. The laser parameters of pulse energy, pulse duration and scan speed remained constant at 1.8 J, 3.0 ms and 350 mm/min, respectively. The gap width and resulting welds are shown in Fig.5.2. An increase in the gap width to 35-55 μm generates a wide void, and further increase of the gap to 55-80 μm produces a wider and larger void. Gap widths of less

Fig. 5.2. Gap width effect on weld appearances at beam location x=0 (Ishak, Yamasaki et al.

Wider gaps over than 80 μm fail to create a joint. A gap wider than about 35-55 μm creates a long crack that starts at root of the two sheets. The crack then expands to the upper surface of the weld bead. Additionally, when the gap width exceeds 35 μm, the amount of molten metal is insufficient, and thus the bond width and penetration depth become small. These results suggest that the gap width is the most important factor in the occurrence of defects as well as in controlling the bond width and penetration depth of thin sheets. Therefore, to minimize the defects and optimize the bond width and penetration depth, the gap width

To analyze the weld geometry of the sample, the average weld width and penetration depth were measured. The effects of the center beam location on the bond width and penetration depth are shown in Figs.5.3 and 5.4. The weld bond width and penetration depth decreased significantly with increasing distance of the beam center location from the edge of the upper specimen. The weld bond width and penetration depth decreased significantly with increasing distance of the beam center from the edge of the upper specimen. A larger

2009)

must be restricted to less than 35 μm.

**5.2.2 Effect of laser beam center position on weld geometry**

than 35 μm significantly reduce the size and create a round shape void.

Figure 5.5 shows the weld microstructure obtained at x=0 with scan speeds of 250, 350 and 450 mm/min. Table 5.1 lists the mean grain size with different scan speeds and locations. Fine grains are apparent at the upper weld with an average grain size of 3.5 μm2 ±2.3 for the 450 mm/min scenario. A mix of fine and medium-large grains with an average grain size of 17 μm2 ±11 can be seen at the middle of weld bead, given a joint between the two sheets. These grain sizes are much smaller than those reported elsewhere, where grain sizes were as large as 20 μm in the fusion zone (Cao, M.Jahazi et al. 2006; Y.Quan 2008).

Fig. 5.3. Bond width at different beam location

Large elongated grains with an average grain size of 141 µm ±89 are apparent in region III, and non-uniform shapes with an average grain size of 13.5 µm ±9.0 can be seen in region IV. The microstructure at location IV is the same as that of the base metal. There are no significant changes in terms of grain size between scan speeds of 350 mm/min and 450 mm/min at each region.

With a scan speed of 250 mm/min, the average grain size in region I is about 27 µm ±18. An increase in the grain size is observed in region II with a grain size of about 53 µm ±31. However, we observed no changes in grain size in region III, while in region IV a larger grain size of about 26 µm ±12 was apparent at the bottom of the weld. The lengths of the long elongated grains decrease as scan speed increases, as listed in table 5.2.

We conclude that the grain size increases with increasing distance from the upper weld because of the changes in thermal cycling during welding. The grain size at the fused area

Laser Welding of Thin Sheet Magnesium Alloys 151

becomes larger at lower scan speeds. A low scan speed will result in impaired cooling. Moreover, since the cooling rate during solidification can be small, the time for grain coarsening will become longer. A large number of precipitated particles will also be apparent in the form of a very fine white deposit in the weld area. It has been reported elsewhere that the sole precipitation phase of AZ31 magnesium alloy is γ-Mg17Al12, and the

white particles have been identified as Mg17(Al, Zn)12 (Quan, Chen et al. 2008).

(Area III and IV)

HAZ length 187 ± 14

**5.2.4 Defects**

**<IP_ADDRESS> Void at the root**

heat transfer to the lower sheet.

in the large grain area in a transgranular manner.

**<IP_ADDRESS> Cracks**

Table 5.2. HAZ length at different scan speed and location

beam location of x=0 and with a scan speed of 400-450 mm/min.

Scan speed , mm/min 250 350 450

Void at the root between the two sheets are apparent at almost all beam center locations and scan speeds. An increase in beam location distance makes the void size larger. Larger voids reduce the bond widths of the joints. By contrast, the size of voids can be reduced by selecting a high scan speed. The void size can be greatly reduced to around 0-40 µm at a

The void or discontinuity at the weld root probably occurred because of a lack of fusion (Yamauchi, Inaba et al. 1982; Nobuyuki, Masahiro et al. 2002; G.Rihar and M.Uran 2006; Z. Barsoum and A. Lundback 2009). This probably happens because of insufficient melting of the lower sheet, although the upper sheet can be heated at the welded root area as shown in Fig. 5.6. This insufficient melting was due to inadequate energy input and weld preparation throughout the welded root area (Yamauchi, Inaba et al. 1982; Nobuyuki, Masahiro et al. 2002; G.Rihar and M.Uran 2006; Z. Barsoum and A. Lundback 2009) . The narrow gap may also contribute to a lack of fusion. Furthermore, we believe that an oxide thin film probably formed on the surface of the lower sheet, although it had been polished using SiC paper of grit 1000 and cleaned using acetone. Magnesium is an active metal, and oxide can easily form on the surface when in contact with air after grinding (Z. Barsoum and A. Lundback 2009). This thin layer may have acted as a barrier, preventing

Cracks are apparent at lower speeds when x=0 and become smaller as the scan speed is increased. No visible cracks were observed at the beam center location x=0 with scan speeds of 400 to 450 mm/min. Furthermore, no macro cracks were observed on the surface of the upper sheet weld bead. The appearance of the cracks proved to be highly consistent across all specimens at low scan speeds. The crack profile indicates that cracking initiated from the root and then propagated into the weld metal but not to the center or upper surface of weld zone. From microstructure images, we conclude that the cracking occurred and propagated

112 ± 8 (Area III) 83 ± 10 Area II

Fig. 5.4. Penetration depth at different beam location

Fig. 5.5. Microstructure of welded joints at scan speed 450 mm/min at four different regions(Ishak, Yamasaki et al. 2009)

Table 5.1. Mean grain size at different scan speed and location

becomes larger at lower scan speeds. A low scan speed will result in impaired cooling. Moreover, since the cooling rate during solidification can be small, the time for grain coarsening will become longer. A large number of precipitated particles will also be apparent in the form of a very fine white deposit in the weld area. It has been reported elsewhere that the sole precipitation phase of AZ31 magnesium alloy is γ-Mg17Al12, and the white particles have been identified as Mg17(Al, Zn)12 (Quan, Chen et al. 2008).

Table 5.2. HAZ length at different scan speed and location

#### **5.2.4 Defects**

150 Nd YAG Laser

**250mm/min 300mm/min 350mm/min 450mm/min**

**0 0.1 0.2**

**Beam center location, -x, mm**

Fig. 5.5. Microstructure of welded joints at scan speed 450 mm/min at four different

Table 5.1. Mean grain size at different scan speed and location

Mean grain size, μm Weld region I II III IV 250 mm/min 27± 18 53 ± 31 98 ± 57 26 ± 12 350 mm/min 4.0 ± 1.5 18 ± 15 160 ± 84 14 ± 10 450 mm/min 3.5 ± 2.0 17 ± 11 141 ± 89 13.5 ± 9.0

Fig. 5.4. Penetration depth at different beam location

**Depth,**

**μ**

**m**

regions(Ishak, Yamasaki et al. 2009)

#### **<IP_ADDRESS> Void at the root**

Void at the root between the two sheets are apparent at almost all beam center locations and scan speeds. An increase in beam location distance makes the void size larger. Larger voids reduce the bond widths of the joints. By contrast, the size of voids can be reduced by selecting a high scan speed. The void size can be greatly reduced to around 0-40 µm at a beam location of x=0 and with a scan speed of 400-450 mm/min.

The void or discontinuity at the weld root probably occurred because of a lack of fusion (Yamauchi, Inaba et al. 1982; Nobuyuki, Masahiro et al. 2002; G.Rihar and M.Uran 2006; Z. Barsoum and A. Lundback 2009). This probably happens because of insufficient melting of the lower sheet, although the upper sheet can be heated at the welded root area as shown in Fig. 5.6. This insufficient melting was due to inadequate energy input and weld preparation throughout the welded root area (Yamauchi, Inaba et al. 1982; Nobuyuki, Masahiro et al. 2002; G.Rihar and M.Uran 2006; Z. Barsoum and A. Lundback 2009) . The narrow gap may also contribute to a lack of fusion. Furthermore, we believe that an oxide thin film probably formed on the surface of the lower sheet, although it had been polished using SiC paper of grit 1000 and cleaned using acetone. Magnesium is an active metal, and oxide can easily form on the surface when in contact with air after grinding (Z. Barsoum and A. Lundback 2009). This thin layer may have acted as a barrier, preventing heat transfer to the lower sheet.

#### **<IP_ADDRESS> Cracks**

Cracks are apparent at lower speeds when x=0 and become smaller as the scan speed is increased. No visible cracks were observed at the beam center location x=0 with scan speeds of 400 to 450 mm/min. Furthermore, no macro cracks were observed on the surface of the upper sheet weld bead. The appearance of the cracks proved to be highly consistent across all specimens at low scan speeds. The crack profile indicates that cracking initiated from the root and then propagated into the weld metal but not to the center or upper surface of weld zone. From microstructure images, we conclude that the cracking occurred and propagated in the large grain area in a transgranular manner.

Laser Welding of Thin Sheet Magnesium Alloys 153

Higher scan speeds can reduce stress as well as leading to the formation of a wider fine equiaxed grain area. Subsequently, the large grain areas become much smaller, as shown in table 5.2. These fine equiaxed grains are less susceptible to cracking than larger ones because

This section present the simulation model of laser lap welding process utilize the efficient solution procedure of the finite element code ANSYS. The physical dimensions of the bond width, penetration depth and HAZ width were measured and compared with the

A moving, pulsed laser beam vertically radiated on the edge of upper sheet. The laser beam center located on edge of upper specimen and laser were scanned towards y direction. The pulse energy is sufficient to form a joint on the specimens. The dimensions of the thin sheets and heat source are shown in Fig. 5.7. The thin sheets with thickness of 0.3 mm is

1. The laser beam is incident vertically towards the surface of the workpiece. The heat or

2. The laser beam spot diameter on surface of specimen is 0.4 mm. Half of the spot beam

3. The workpiece is initially at 20 °C. Both the laser beam and the coordinate mesh are fixed and the work pieces moves in the positive y-direction with a constant velocity, v.

5. The absorption of laser is fixed at 18 %. This based on the experimental results which

4 mm

The following assumptions were made in the formulation of the finite element model:

laser beam center is located at edge of upper specimen.

56 mm

hit on upper sheet and other half on lower sheet.

4. All thermophysical are considered to be constant.

Fig. 5.7. Schematic drawing 3D geometrical model

the stress is more evenly distributed among numerous grain boundaries.

**5.3 Numerical simulation of lap fillet welding**

**5.3.1 Geometrical model description**

prediction model.

overlapped each other.

are around 18-20 %.

Fig. 5.6. Microstructure around void at root

(W.Zhou, Long et al. 2007)reported that a crack was observed during tungsten inert gas welding of AZ91D related to the liquation of the second phase or low melting point precipitates in the heat affected zone (HAZ) or partial melting zone (PMZ). This type of crack is known as "liquation cracking" due to the existence of a low melting point for intermetallic compounds at grain boundaries of the HAZ and PMZ. These effects can greatly decrease the strength of the weld (W.Zhou, Long et al. 2007). However based on EDS analysis there was no segregation of Zn, Al and Mn in that area could be detected. Secondary phases of low melting intermetallic compounds consisting of Mg and Al or Zn did not precipitate near the crack areas. No finer grains were apparent at the open crack, which was likely observed in reference (W.Zhou, Long et al. 2007). Furthermore, no significant hardness was detected in this area. This further shows that precipitation of second phase low melting intermetallic compounds did not take place in the crack region.

However, a peak of O was apparent near the crack's open surface. Although the surface of the thin sheet had been polished with SiC paper prior to welding, we believe that a thin oxide film probably still formed because magnesium is a very reactive metal. The crack probably initiated from the void at the root due to high stress concentrations, then propagated to the large elongated grain area along the thin oxide film.

During low scan speed welding, a wider area of large grains probably formed at the boundary of the fusion zone and base metal. This area was exposed to much lower temperature than the middle of the fusion zone; thus, oxide films were not properly broken or melted during welding. Furthermore, a low scan speed may result in high heat input, which consequently leads to higher stresses in the large grain area as the metal solidifies and contracts. Therefore, cracks could easily propagate throughout the oxide film area.

**grain area**

**Not Melt**

**area Large size**

**Fine/mix size grain**

(W.Zhou, Long et al. 2007)reported that a crack was observed during tungsten inert gas welding of AZ91D related to the liquation of the second phase or low melting point precipitates in the heat affected zone (HAZ) or partial melting zone (PMZ). This type of crack is known as "liquation cracking" due to the existence of a low melting point for intermetallic compounds at grain boundaries of the HAZ and PMZ. These effects can greatly decrease the strength of the weld (W.Zhou, Long et al. 2007). However based on EDS analysis there was no segregation of Zn, Al and Mn in that area could be detected. Secondary phases of low melting intermetallic compounds consisting of Mg and Al or Zn did not precipitate near the crack areas. No finer grains were apparent at the open crack, which was likely observed in reference (W.Zhou, Long et al. 2007). Furthermore, no significant hardness was detected in this area. This further shows that precipitation of second phase low melting intermetallic compounds did not take place in the crack region.

However, a peak of O was apparent near the crack's open surface. Although the surface of the thin sheet had been polished with SiC paper prior to welding, we believe that a thin oxide film probably still formed because magnesium is a very reactive metal. The crack probably initiated from the void at the root due to high stress concentrations, then

During low scan speed welding, a wider area of large grains probably formed at the boundary of the fusion zone and base metal. This area was exposed to much lower temperature than the middle of the fusion zone; thus, oxide films were not properly broken or melted during welding. Furthermore, a low scan speed may result in high heat input, which consequently leads to higher stresses in the large grain area as the metal solidifies and

contracts. Therefore, cracks could easily propagate throughout the oxide film area.

propagated to the large elongated grain area along the thin oxide film.

Fig. 5.6. Microstructure around void at root

Higher scan speeds can reduce stress as well as leading to the formation of a wider fine equiaxed grain area. Subsequently, the large grain areas become much smaller, as shown in table 5.2. These fine equiaxed grains are less susceptible to cracking than larger ones because the stress is more evenly distributed among numerous grain boundaries.

#### **5.3 Numerical simulation of lap fillet welding**

This section present the simulation model of laser lap welding process utilize the efficient solution procedure of the finite element code ANSYS. The physical dimensions of the bond width, penetration depth and HAZ width were measured and compared with the prediction model.

#### **5.3.1 Geometrical model description**

A moving, pulsed laser beam vertically radiated on the edge of upper sheet. The laser beam center located on edge of upper specimen and laser were scanned towards y direction. The pulse energy is sufficient to form a joint on the specimens. The dimensions of the thin sheets and heat source are shown in Fig. 5.7. The thin sheets with thickness of 0.3 mm is overlapped each other.

The following assumptions were made in the formulation of the finite element model:

Fig. 5.7. Schematic drawing 3D geometrical model

Laser Welding of Thin Sheet Magnesium Alloys 155

Laser heat input was on the surface as heat flux. Heat losses due to convection and radiation are very small (Yeung and P.H 1999) and hence also there were not considered in this analysis, as they make calculation non-linear, thus causing convergence problems and

The finite element code ANSYS provides the convenient means of numerically modeling pulsed laser welding. The solution technique is chosen depending on the type of problem. In the present case, the thermal history of lap fillet weld is required, so a transient thermal analysis must be performed. This requires an integration of the heat conduction equation with respect to time. In the finite element formulation, this equation can be written for each

[ ( ){ } [ ( ){ } { } { ( )} *CT T KT T V QT* (5.2)

, , *xyz kkk* = thermal conductivity in the x, y, and z directions (W/mK)

*c* = specific heat capacity (J/kgK)

*v* = velocity of workpiece (m/s)

increasing computing time by a great amount.

element as follow (Frewin and Scott 1999);

[ ] *K* = conductivity matrix [ ] *C* = specific heat matrix

{Q} = nodal heat flow vector

{T} = vector or nodal temperature { } *T* = vector of time derivative of {T}

**5.3.3 Numerical result and discussion**

{ } *V* = velocity vector for the moving workpiece

applied on the thin sheet surface at the edge of upper sheets.

This equation is simply the vector and matrix equivalent of Equation 4.4.

The predicted temperature distributions from numerical analysis at different scan speeds are shown in through Fig. 5.9. It is noted that the temperature distribution has a parabolic form, which is the consequence of the Gaussian energy profile. The maximum temperature is recorded at the central axis of laser beam. It is explained by the fact that the heat source is

The maximum temperature decreases as scan speed increases as a result of lower heat input. At scan speed of 250 mm/min the melt area are larger than scan speeds of 350 mm/min and 450 mm/min. It is also noted that the area of temperature between 500 °C until 575 °C becomes smaller as scan speed increases. This result because of heat input decreases as scan speed increases. From the results, it shows that the weld area slightly larger at starting (starting calculation time) and end point (end calculation time) of specimen compare with at the middle of the specimen. It shows similar tendency with the experimental results where wider weld zone and deeper penetration obtained at edge compare with at the middle of the specimen. This occurs because of lower heat concentration at middle than at the edge of

= density (kg/m3)

*t* = time (s)

where

The mesh structure of model is shown in Fig. 5.8. The mesh size is divided into 5 different zones. As the distance of specimens far away from weld area, the mesh becomes rougher. The finest mesh is defined at weld area where 84 mesh along the 4 mm width of specimen. Free mesh is defined at zone 1, while mapped mesh at zone 2, 3, 4 and 5. The thickness of mesh at zone 2 to zone 4 consists of 7 layers of one thin sheet, whereas, zone 1 only 1 layer. In this simulation, the pulse energy is fixed at 1.8 J, pulse duration 3.0 ms, repetition rate at 80 Hz. The simulation is done at three different scan speeds 250 mm/min, 350 mm/min and 450 mm/min. The simulation time was 0.53 s, 0.375 s and 0.2975 s for scan speed 250 mm/min, 350 mm/min and 450 mm/min, respectively.

Fig. 5.8. Mesh structure at weld area

#### **5.3.2 Numerical description**

The temperature field caused by moving laser beam is transient in nature. Here, in order to simulate the process, a three-dimensional transient heat transfer model was constructed. The general three dimensional heat diffusion equations is of the form

$$\frac{\partial}{\partial \mathbf{x}} \left( k\_x \frac{\partial T}{\partial \mathbf{x}} \right) + \frac{\partial}{\partial y} \left( k\_y \frac{\partial T}{\partial y} \right) + \frac{\partial}{\partial z} \left( k\_z \frac{\partial T}{\partial z} \right) + Q = \rho \mathbf{c} \left( \hat{\sigma} \frac{\partial T}{\partial t} - v \frac{\partial T}{\partial \mathbf{x}} \right) \tag{5.1}$$

where,

 *xyz* , , = coordinate system attached to the heat source *Q* = power generation per unit volume (W/m3)

, , *xyz kkk* = thermal conductivity in the x, y, and z directions (W/mK)

*c* = specific heat capacity (J/kgK)

154 Nd YAG Laser

6. The gap between the thin sheets is 30 µm and is divided into two layers which thickness

7. As pulsed laser welding involves very rapid melting and solidification, convective redistribution of heat within the weld pool is not a significant as it is in other processes

The mesh structure of model is shown in Fig. 5.8. The mesh size is divided into 5 different zones. As the distance of specimens far away from weld area, the mesh becomes rougher. The finest mesh is defined at weld area where 84 mesh along the 4 mm width of specimen. Free mesh is defined at zone 1, while mapped mesh at zone 2, 3, 4 and 5. The thickness of mesh at zone 2 to zone 4 consists of 7 layers of one thin sheet, whereas, zone 1 only 1 layer. In this simulation, the pulse energy is fixed at 1.8 J, pulse duration 3.0 ms, repetition rate at 80 Hz. The simulation is done at three different scan speeds 250 mm/min, 350 mm/min and 450 mm/min. The simulation time was 0.53 s, 0.375 s and 0.2975 s for scan speed 250

The temperature field caused by moving laser beam is transient in nature. Here, in order to simulate the process, a three-dimensional transient heat transfer model was constructed.

> *T T T TT k k k Qc v x x y y zz t x*

(5.1)

The general three dimensional heat diffusion equations is of the form

*xyz*

 *xyz* , , = coordinate system attached to the heat source *Q* = power generation per unit volume (W/m3)

where liquid pool is permanent. Convective flow of heat, therefore, is neglected.

mm/min, 350 mm/min and 450 mm/min, respectively.

Gap

Fig. 5.8. Mesh structure at weld area

Upper sheet

Lower sheet

**5.3.2 Numerical description**

where,

of 15 µm each.

*v* = velocity of workpiece (m/s)

Laser heat input was on the surface as heat flux. Heat losses due to convection and radiation are very small (Yeung and P.H 1999) and hence also there were not considered in this analysis, as they make calculation non-linear, thus causing convergence problems and increasing computing time by a great amount.

The finite element code ANSYS provides the convenient means of numerically modeling pulsed laser welding. The solution technique is chosen depending on the type of problem. In the present case, the thermal history of lap fillet weld is required, so a transient thermal analysis must be performed. This requires an integration of the heat conduction equation with respect to time. In the finite element formulation, this equation can be written for each element as follow (Frewin and Scott 1999);

$$\{\mathbf{C}(T)\{\dot{T}\} + \{\mathbf{K}(T)\{T\} + \{V\} = \{Q(T)\}\}\tag{5.2}$$

where

This equation is simply the vector and matrix equivalent of Equation 4.4.

#### **5.3.3 Numerical result and discussion**

The predicted temperature distributions from numerical analysis at different scan speeds are shown in through Fig. 5.9. It is noted that the temperature distribution has a parabolic form, which is the consequence of the Gaussian energy profile. The maximum temperature is recorded at the central axis of laser beam. It is explained by the fact that the heat source is applied on the thin sheet surface at the edge of upper sheets.

The maximum temperature decreases as scan speed increases as a result of lower heat input. At scan speed of 250 mm/min the melt area are larger than scan speeds of 350 mm/min and 450 mm/min. It is also noted that the area of temperature between 500 °C until 575 °C becomes smaller as scan speed increases. This result because of heat input decreases as scan speed increases. From the results, it shows that the weld area slightly larger at starting (starting calculation time) and end point (end calculation time) of specimen compare with at the middle of the specimen. It shows similar tendency with the experimental results where wider weld zone and deeper penetration obtained at edge compare with at the middle of the specimen. This occurs because of lower heat concentration at middle than at the edge of

Laser Welding of Thin Sheet Magnesium Alloys 157

Experiment width Experiment depth Experiment HAZ Numerical width Numerical depth Numerical HAZ

Fig. 5.11. Comparison of weld geometry and HAZ width between experimental and

We investigated the influences of welding parameters, such as beam center location, laser scan speed and gap size, on weldability in the context of lap fillet welding of AZ31B magnesium alloys with a thickness of 0.3 mm with a pulsed Nd:YAG laser. The following

**200 250 300 350 400 450 500 550 600**

**Scan speed, mm/min**

For bead on plate (BOP) welding undercut, cracks and pores can be greatly reduced when the process parameters were properly chosen to weld bead on the AZ31B magnesium alloy sheet in the Nd:YAG laser welding system. The depth of penetration decreases as focus point located on surface of thin sheet magnesium alloys. The defects such as undercut, cracks and pores are significantly reduced as pulse energy at 1.8 J, repetition rate at 80 Hz,

The gap width exerts a major influence on the occurrence of joints and defects, such as cracks and voids, between the two sheets. A gap width of less than 35 μm significantly reduces the number of cracks and voids and also increases the penetration depth and bond

Voids and discontinuities at the root occurred because of lack of fusion. A narrow gap oxide layer on the lower surface acted as a heat barrier to prevent melting of the lower sheet. High scan speeds significantly improved the number of voids and discontinuities at the root. Cracks at the weld originated from voids at the root, which propagated in a transgranular manner within the large grain area. No segregation of Al, Mn and Zn occurred at the crack

pulse duration 3.0 ms and scan speed range from 300 to 500 mm/min.

numerical results at different scan speeds

conclusions can be drawn from this work:

**6. Conclusion**

**Width and depth,**

**μ**

**m**

width.

specimen. It is due to heat easily dissipated away as a result of high conductivity at the middle of specimen.

Fig. 5.9. Result of numerical simulation at scan speed 450 mm/min at t=0.2975s

Fig. 5.10 shows the comparison between the numerical simulation result and cross section of experimental results. It shows that at higher scan speeds, the weld geometry between simulation and the experimental results are almost similar. Fig. 5.11 shows quantitative comparison between numerical and experimental results of fusion zone width, penetration depth and HAZ width at weld root. The fusion zone area is the area recorded temperature from melting point. Meanwhile, the HAZ width is measured at weld root with temperature between 500° C until 575 °C. The range of HAZ temperature is based on grain growth studies of AZ31B (Mehtedi, Balloni et al. ; Ben-Artzy, Shtechman et al. 2000; Maryaa, Hector et al. 2006; Sun, Wu et al. 2009). The results of comparison between numerical calculation and experimental results show a reasonable agreement. At lowest scan speed of 250 mm/min, numerical result shows larger fusion width compare with experimental result. The different occurs because of larger defects (void and crack) at the weld root, and occurrence of severe undercut which result in smaller weld width of experimental result. The defects improve as scan speed increases. As a result, good agreement can be obtained at scan speed of 350 mm/min and 450 mm/min.

Fig. 5.10. Comparison between experimental and numerical result at scan speed 250 mm/min

specimen. It is due to heat easily dissipated away as a result of high conductivity at the

Fig. 5.9. Result of numerical simulation at scan speed 450 mm/min at t=0.2975s

Fig. 5.10. Comparison between experimental and numerical result at scan speed 250

scan speed of 350 mm/min and 450 mm/min.

mm/min

Fig. 5.10 shows the comparison between the numerical simulation result and cross section of experimental results. It shows that at higher scan speeds, the weld geometry between simulation and the experimental results are almost similar. Fig. 5.11 shows quantitative comparison between numerical and experimental results of fusion zone width, penetration depth and HAZ width at weld root. The fusion zone area is the area recorded temperature from melting point. Meanwhile, the HAZ width is measured at weld root with temperature between 500° C until 575 °C. The range of HAZ temperature is based on grain growth studies of AZ31B (Mehtedi, Balloni et al. ; Ben-Artzy, Shtechman et al. 2000; Maryaa, Hector et al. 2006; Sun, Wu et al. 2009). The results of comparison between numerical calculation and experimental results show a reasonable agreement. At lowest scan speed of 250 mm/min, numerical result shows larger fusion width compare with experimental result. The different occurs because of larger defects (void and crack) at the weld root, and occurrence of severe undercut which result in smaller weld width of experimental result. The defects improve as scan speed increases. As a result, good agreement can be obtained at

middle of specimen.

Fig. 5.11. Comparison of weld geometry and HAZ width between experimental and numerical results at different scan speeds

#### **6. Conclusion**

We investigated the influences of welding parameters, such as beam center location, laser scan speed and gap size, on weldability in the context of lap fillet welding of AZ31B magnesium alloys with a thickness of 0.3 mm with a pulsed Nd:YAG laser. The following conclusions can be drawn from this work:

For bead on plate (BOP) welding undercut, cracks and pores can be greatly reduced when the process parameters were properly chosen to weld bead on the AZ31B magnesium alloy sheet in the Nd:YAG laser welding system. The depth of penetration decreases as focus point located on surface of thin sheet magnesium alloys. The defects such as undercut, cracks and pores are significantly reduced as pulse energy at 1.8 J, repetition rate at 80 Hz, pulse duration 3.0 ms and scan speed range from 300 to 500 mm/min.

The gap width exerts a major influence on the occurrence of joints and defects, such as cracks and voids, between the two sheets. A gap width of less than 35 μm significantly reduces the number of cracks and voids and also increases the penetration depth and bond width.

Voids and discontinuities at the root occurred because of lack of fusion. A narrow gap oxide layer on the lower surface acted as a heat barrier to prevent melting of the lower sheet. High scan speeds significantly improved the number of voids and discontinuities at the root. Cracks at the weld originated from voids at the root, which propagated in a transgranular manner within the large grain area. No segregation of Al, Mn and Zn occurred at the crack

Laser Welding of Thin Sheet Magnesium Alloys 159

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Moon Jonghyun, K. S., Mizutani Masami, and Matsunawa Akira (2002). "Lap welding

Nobuyuki, A., T. Masahiro, et al. (2002). "Welding of Aluminum Alloy with High Power

Quan, Y. J., Z. H. Chen, et al. (2008). "Effects of heat input on microstructure and tensile

Sun, P.-h., H.-y. Wu, et al. (2009). Deformation Characteristics of Fine-grained Magnesium

Toshikatsu Asahina, H. T., Hitaka Itoh and Shigekazu Taguchi (2005). Some characteristics

Ukita, S., T. Akamatsu, et al. (1993). "The welding conditions of very thin aluminum sheet of

W.Zhou, T. Z. Long, et al. (2007). "Hot cracking in tungsten inert gas welding of magnesium

Y.Quan, S. C., Z. Yu, X.S. Gong, M.Li, (2008). "Characteristics of laser welded wrought Mg-

Yamauchi, N., Y. Inaba, et al. (1982). "Formation mechanism of lack of fusion in MAG

Yeung, K. S. and T. P.H (1999). "Transient thermal analysis of spot welding electrodes"

Z. Barsoum and A. Lundback (2009). "Simplified FE welding simulation of fillet welds-3D effects on the formation residual stresses." Eng. Fail. Anal. 16(7): 2281-2289. Zhao, H. and T. Debroy (2001). "Pore Formation During Laser Beam Welding Of Die-Cast

Magnesium Alloy AM60B —Mechanism And Remedy." Weld. Research Supp:

3(9): 1045-1056.

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204s-210s.

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alloy." Materials letters 60: 2194-2197.

(In Japanese)." Japan welding society 20: 468-476.

Direct Diode Laser, ." Trans. JWRI 31(2): 157-163.

high welding." Japan welding society 11: 361-364.

alloy AZ91D." Mat. Sci. and Tech. 23 (11): 1294-1299.

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with pulsed Nd:YAG laser." Journal of Solid Mechanics and Materials Engineering

its tensile deformation and failure behaviors." Materials Science and Engineering,

ZM21 and AZ31 magnesium alloys." Metallurgical science and technology: 23-

characteristics of thin sheet metals with combined laser beams of different length

properties of laser welded magnesium alloy AZ31." Mater. Charact. 59(10): 1491-

Alloy AZ31B Thin Sheet during Fast Gas Blow Forming,. Proceedings of the International MultiConference of Engineers and Computer Scientists 2009 (IMECS

of pulsed YAG laser welds of Magnesium Alloys. Nihon University research report

opening, whereas a high content of oxide may have caused the crack/discontinuity. Higher scan speeds significantly reduced the number of defects because of the narrower large grain area and the wider fine grain area.

Lap fillet welding of the thin sheet produced a weld bead width of less than 1 mm, with no melt-through and good weld bead appearance, when the beam center location was on the edge of the upper specimen and when we used scan speeds of 350 to 450 mm/min. Under the optimized conditions, a large fine grain size area can be obtained from the top to the middle of the weld bead, and the HAZ becomes very narrow.

The weld geometry (weld width, depth and HAZ width) of numerical simulation results show reasonable agreement with experimental results at high scan speed of 350 mm/min and 450 mm/min.

#### **7. References**

opening, whereas a high content of oxide may have caused the crack/discontinuity. Higher scan speeds significantly reduced the number of defects because of the narrower large grain

Lap fillet welding of the thin sheet produced a weld bead width of less than 1 mm, with no melt-through and good weld bead appearance, when the beam center location was on the edge of the upper specimen and when we used scan speeds of 350 to 450 mm/min. Under the optimized conditions, a large fine grain size area can be obtained from the top to the

The weld geometry (weld width, depth and HAZ width) of numerical simulation results show reasonable agreement with experimental results at high scan speed of 350 mm/min

Aghios, E., B. Bronfiu, et al. (2001). "The role of the magnesium industry in protecting the environment." Journal of Materials Processing Technology 117: 381-385. Ben-Artzy, A., A. Shtechman, et al. (2000). "Deformation characteristics of wrought

Cao, X., M. Jahazi, et al. (2006). " A review of laser welding techniques for magnesium

Cao, X., M.Jahazi, et al. (2006). "A review of laser welding techniques for magnesium alloys."

Costaa, A. P., L. s. Quintinoa, et al. (2003). " Laser beam welding hard metals to steel."

Czerwinski, F. (2002). "The oxidation behavior of an AZ91D magnesium alloy at high

Frewin, M. R. and D. A. Scott (1999). "Finite element model of pulsed laser welding."

G. Ben-Hamu, D. E., and C. E Cross, Th. Bollinghaus, , . (2007). "The relation between

G.Rihar and M.Uran (2006). "Lack of fusion-Characterisation of indications." Welding in the

Haferkamp, H., A. Ostendorf, et al. (2003). Laser beam welding of magnesium alloys a

Ishak, M. (2010). Study on Lap Fillet Welding of AZ31B Magnesium Thin Sheets using

microstructure and corrosion behavior of GTA welded AZ31B magnesium sheet."

process at the threshold to the industrial manufacturing. International Conference

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magnesium alloys AZ31, ZK60." Magnesium Technology: 363-374.

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Ion, J. C. (2005). Laser Processing of Engineering Materials, Elsevier.

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Duley, W. W. (1999). Laser Welding, John Wiley & Sons, Inc.

Welding research supplement: 15s-21s.

world issue ½ 50(1/2): 35-39.

University, Japan. Phd thesis.

area and the wider fine grain area.

and 450 mm/min.

**7. References**

middle of the weld bead, and the HAZ becomes very narrow.

**Micro-Welding of Super Thermal Conductive**

Mohd Idris Shah Ismail1,2, Yasuhiro Okamoto1 and Akira Okada1 *1Graduate School of Natural Science and Technology, Okayama University*

*2Department of Mechanical & Manufacturing Engineering, Universiti Putra Malaysia*

The diffusion of generated heat in the electronic devices is an important issue. The heat would be diffused from electronic devices by passive strategies, which would be carried out by the use of high thermal conductivity materials as a heat sink. The development of advanced materials with the superior high-thermal properties and the high strength-toweight ratio has led to new metal matrix composites (MMCs) as a great attractive material in the electrical and electronic industries. Aluminum and its alloys are widely used for the manufacturing of MMCs, which have reached the industrial stage in some areas (Barekar et al., 2009). In order to manufacture practical components from MMCs, a technique for joining MMCs to other similar composites or to monolithic materials is strongly required. Therefore, the development of reliable and economic joining technique is important for extending the applications of MMCs. It is well-known that laser welding is the most flexible and versatile welding technology, and it has succeeded in the welding of MMCs (Niu et al.,

Recently, a super thermal conductive (STC) aluminum-graphite (Al-Gr) composite with the high thermal conductivity and the low coefficient of thermal expansion (CTE) was developed (Ueno et al., 2009). The properties of thermal conductivity versus thermal expansion coefficient are summarized in Figure 1, where the upper ellipse zone shows the STC composite materials, and the other of conventional thermal conductive materials such as Cu, Al, Al-SiC, Cu-W, Cu-Mo, AlN and Si are also indicated. It is difficult to weld STC Al-Gr composite material compared with existing MMCs, since the graphite material only can be melted under the high pressure with the high temperature (Kirillin & Kostanovskii, 2003). In the laser welding, the use of standard pulse profile is limited to joint this material, since the uncontrolled heat input generates an overshoot, which leads to undesirable welded joints. The control of heat input by a pulse waveform is very important to achieve a suitable welding penetration, preventions of overheating and

In this study, the welding of STC Al-Gr composite was experimentally and numerically investigated by using a pulsed Nd:YAG laser with the control of pulse waveforms, which

**1. Introduction**

2006; Bassani et al., 2007).

unacceptable welding defects.

**Composite by Pulsed Nd:YAG Laser**

*1Japan 2Malaysia*

Zhu, J., L. Li, et al. (2005). "CO2 and diode laser welding of AZ31 magnesium alloy." Applied Surface Science 247 pp. 300-306. **9**

## **Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser**

Mohd Idris Shah Ismail1,2, Yasuhiro Okamoto1 and Akira Okada1 *1Graduate School of Natural Science and Technology, Okayama University 2Department of Mechanical & Manufacturing Engineering, Universiti Putra Malaysia 1Japan 2Malaysia*

#### **1. Introduction**

160 Nd YAG Laser

Zhu, J., L. Li, et al. (2005). "CO2 and diode laser welding of AZ31 magnesium alloy."

The diffusion of generated heat in the electronic devices is an important issue. The heat would be diffused from electronic devices by passive strategies, which would be carried out by the use of high thermal conductivity materials as a heat sink. The development of advanced materials with the superior high-thermal properties and the high strength-toweight ratio has led to new metal matrix composites (MMCs) as a great attractive material in the electrical and electronic industries. Aluminum and its alloys are widely used for the manufacturing of MMCs, which have reached the industrial stage in some areas (Barekar et al., 2009). In order to manufacture practical components from MMCs, a technique for joining MMCs to other similar composites or to monolithic materials is strongly required. Therefore, the development of reliable and economic joining technique is important for extending the applications of MMCs. It is well-known that laser welding is the most flexible and versatile welding technology, and it has succeeded in the welding of MMCs (Niu et al., 2006; Bassani et al., 2007).

Recently, a super thermal conductive (STC) aluminum-graphite (Al-Gr) composite with the high thermal conductivity and the low coefficient of thermal expansion (CTE) was developed (Ueno et al., 2009). The properties of thermal conductivity versus thermal expansion coefficient are summarized in Figure 1, where the upper ellipse zone shows the STC composite materials, and the other of conventional thermal conductive materials such as Cu, Al, Al-SiC, Cu-W, Cu-Mo, AlN and Si are also indicated. It is difficult to weld STC Al-Gr composite material compared with existing MMCs, since the graphite material only can be melted under the high pressure with the high temperature (Kirillin & Kostanovskii, 2003). In the laser welding, the use of standard pulse profile is limited to joint this material, since the uncontrolled heat input generates an overshoot, which leads to undesirable welded joints. The control of heat input by a pulse waveform is very important to achieve a suitable welding penetration, preventions of overheating and unacceptable welding defects.

In this study, the welding of STC Al-Gr composite was experimentally and numerically investigated by using a pulsed Nd:YAG laser with the control of pulse waveforms, which

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 163

phase diagram (Massalski et al., 1990), it can be seen that the coexistence of liquid aluminum and solid carbon requires the existence of an intermediate phase Al4C3. At equilibrium phase, it is impossible to directly dissolve solid carbon into liquid aluminum unless the temperature is above the melting point of Al4C3 (Lu et al., 2001). Graphite, in the form of fibers or particulates has been recognized as a high-strength and low density material (Barekar et al., 2009). The material combination of aluminum and graphite could produce an advanced aluminum-graphite (Al-Gr) composite, which has an unique thermal properties,

Recently, a super thermal conductive(STC) Al-Gr composite was developed, which have been prepared by utilizing a pulsed electric current sintering method (Ueno et al., 2009) to realize the properties of higher thermal conductivity and lower thermal expansion coefficient. The material properties of STC Al-Gr composite are shown in Table 1. It shows that the thermal conductivity, thermal expansion coefficient and bending strength are anisotropic due to the directional properties of graphite and their designed orientation in xy plane-direction (top view) and xz thickness-direction (section view) as

The demand for lighter and thinner products has led MMCs to be manufactured, especially in welding process. The welding process of MMCs has attracted various industries by the reduction of material weight and costs, and improvement of design flexibility. Therefore, the development of efficient, flexible and reliable joining technique is significantly important to increase the potential of MMCs. Compared with the conventional fusion welding techniques, laser welding has numerous advantages such as narrow heat-affected zone (HAZ), good quality of weld bead, precise positioning and control of irradiation beam and its movement, the ease of automation and the high production speed (Behler et al., 1997; Pan et al., 2004). Laser welding is characterized by intermittent laser beam powers that would

due to to the opposite thermal expansion coefficients of aluminum and graphite.

Fig. 2. Al-C phase diagram

shown in Figure 3.

can provide a well-controlled heat input with high energy density. These investigations have led to an optimum welding condition proposed for a pulsed laser welding with minimum weld defects. The experimental work was carried out in two sections, namely the bead-on-plate welding of STC Al-Gr composite, and the overlap welding of pure aluminum and STC Al-Gr composite. The welding characteristics of STC Al-Gr composite were discussed by the observation of joint part with optical microscope, scanning electron microscope (SEM) and energy dispersive spectroscopy (EDS). Moreover, the temperature distributions in the laser micro-welding process were numerically analyzed to discuss the proper heat input. The weld strength was also evaluated by a shearing test for the overlap welding with and without the control of pulse waveform.

Fig. 1. Thermal conductivity and thermal expansion coefficient of STC composites and other materials

#### **2. Background and overview**

Each material has its own unique properties. In many applications, it has been found that there is a need for the combination of several properties together. These desired material property combinations can be achieved by the development of new composite materials. Composite materials are usually classified by the type of material used for the matrix. The four primary categories of composites are polymer matrix composites (PMCs), metal matrix composites (MMCs), ceramic matrix composites (CMCs), and carbon-carbon composites (CCCs) (Kutz, 2002). In this study, the viewpoint was MMCs, which having the properties of lightweight, high specific strength, wear resistance, and a low coefficient of thermal expansion.

Aluminum has many properties such as high thermal and electrical conductivity, high corrosion resistance, low cost and lightweight that make it ideal for use in MMCs. In addition, aluminum can be recycled and offers intriguing environmental and economical opportunities (Cannillo et al., 2010). From the chemical point of view, the aluminum reacts with carbon to form only aluminum carbide (Al4C3) intermetallic compounds as shown in the Figure 2. The Al4C3 is known as the only stable intermediate in the Al-C binary system. However, it is very brittle at the ambient temperature (Lu et al., 2001). Based on the Al-C phase diagram (Massalski et al., 1990), it can be seen that the coexistence of liquid aluminum and solid carbon requires the existence of an intermediate phase Al4C3. At equilibrium phase, it is impossible to directly dissolve solid carbon into liquid aluminum unless the temperature is above the melting point of Al4C3 (Lu et al., 2001). Graphite, in the form of fibers or particulates has been recognized as a high-strength and low density material (Barekar et al., 2009). The material combination of aluminum and graphite could produce an advanced aluminum-graphite (Al-Gr) composite, which has an unique thermal properties, due to to the opposite thermal expansion coefficients of aluminum and graphite.

Fig. 2. Al-C phase diagram

162 Nd YAG Laser

can provide a well-controlled heat input with high energy density. These investigations have led to an optimum welding condition proposed for a pulsed laser welding with minimum weld defects. The experimental work was carried out in two sections, namely the bead-on-plate welding of STC Al-Gr composite, and the overlap welding of pure aluminum and STC Al-Gr composite. The welding characteristics of STC Al-Gr composite were discussed by the observation of joint part with optical microscope, scanning electron microscope (SEM) and energy dispersive spectroscopy (EDS). Moreover, the temperature distributions in the laser micro-welding process were numerically analyzed to discuss the proper heat input. The weld strength was also evaluated by a shearing test for the overlap

Fig. 1. Thermal conductivity and thermal expansion coefficient of STC composites and other

Each material has its own unique properties. In many applications, it has been found that there is a need for the combination of several properties together. These desired material property combinations can be achieved by the development of new composite materials. Composite materials are usually classified by the type of material used for the matrix. The four primary categories of composites are polymer matrix composites (PMCs), metal matrix composites (MMCs), ceramic matrix composites (CMCs), and carbon-carbon composites (CCCs) (Kutz, 2002). In this study, the viewpoint was MMCs, which having the properties of lightweight,

Aluminum has many properties such as high thermal and electrical conductivity, high corrosion resistance, low cost and lightweight that make it ideal for use in MMCs. In addition, aluminum can be recycled and offers intriguing environmental and economical opportunities (Cannillo et al., 2010). From the chemical point of view, the aluminum reacts with carbon to form only aluminum carbide (Al4C3) intermetallic compounds as shown in the Figure 2. The Al4C3 is known as the only stable intermediate in the Al-C binary system. However, it is very brittle at the ambient temperature (Lu et al., 2001). Based on the Al-C

high specific strength, wear resistance, and a low coefficient of thermal expansion.

welding with and without the control of pulse waveform.

materials

**2. Background and overview**

Recently, a super thermal conductive(STC) Al-Gr composite was developed, which have been prepared by utilizing a pulsed electric current sintering method (Ueno et al., 2009) to realize the properties of higher thermal conductivity and lower thermal expansion coefficient. The material properties of STC Al-Gr composite are shown in Table 1. It shows that the thermal conductivity, thermal expansion coefficient and bending strength are anisotropic due to the directional properties of graphite and their designed orientation in xy plane-direction (top view) and xz thickness-direction (section view) as shown in Figure 3.

The demand for lighter and thinner products has led MMCs to be manufactured, especially in welding process. The welding process of MMCs has attracted various industries by the reduction of material weight and costs, and improvement of design flexibility. Therefore, the development of efficient, flexible and reliable joining technique is significantly important to increase the potential of MMCs. Compared with the conventional fusion welding techniques, laser welding has numerous advantages such as narrow heat-affected zone (HAZ), good quality of weld bead, precise positioning and control of irradiation beam and its movement, the ease of automation and the high production speed (Behler et al., 1997; Pan et al., 2004). Laser welding is characterized by intermittent laser beam powers that would

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 165

In addition, the Al4C3 formation reduces the toughness of the weld bead, even it results in the increase of hardness (Bassani et al., 2007). Therefore, preventing the formation of the Al4C3 during welding process is important for successful welding of Al-SiC composites

There are severe heterogeneity of the material structure and the great difference in both physical and chemical properties between the aluminum matrix and graphite particles. This point would make the welding of STC Al-Gr composites to a challenging technological problem for applying the laser welding technique in the industrial application. In this study, the pulsed Nd:YAG laser was used as a laser source. Square shape pulses are the standard output of this laser with the constant power distribution during its duration time. By utilizing the ability of pulsed Nd:YAG laser to shape the temporal power profile of each pulse offers the high flexibility in optimising the weld parameters. This pulse waveform allows the control of penetration, melt pool geometry and keyhole formation. It also has been reported that pulse waveform to be an effective method to reduce or eliminate the

A schematic diagram of experimental setup is shown in Figure 4. In this study, a pulsed Nd:YAG laser (LASAG SLS200 CL8) of 1064 nm in wavelength was used as a laser source. The laser beam was collimated to 15 mm in diameter and delivered by an optical fiber of 50 μm core diameter. The collimator was installed between the optical fiber and the bending mirror, and the collimated laser beam was focused on the specimen surface by a lens of 50 mm in focal length. In order to avoid the back-reflection of incident laser beam, the processing head was aligned by 10 degrees to the perpendicular axis of the specimen surface. The irradiation experiments were carried out in two sections, namely the bead-onplate welding of STC Al-Gr composite, and the overlap welding of pure aluminum and STC Al-Gr composite. The welding experiments were done in a shielding gas of nitrogen with 13L/min flow rate. The stage controller could determine the movement of X-Y-Z stage, and

(Wang et al., 2000).

weld defects (Zhang et al., 2008).

**3. Experimental procedures**

it also synchronized the laser pulse.

Fig. 4. Schematic diagram of experimental setup

allow melting and solidification consecutively. The great advantage of laser welding is keyhole effect by very high power densities involved in the laser welding, which makes laser welding technique widely used in many industrial productions.

Table 1. Material properties of STC Al-Gr composite

Fig. 3. Schematic illustration and optical micrographs of STC Al-Gr composite

Since the STC Al-Gr composite is a new composite in MMCs, no work has been reported concerning the welding behaviour on this composite by laser welding technique. There are several published papers related to the laser welding of MMCs, especially in the composites with combination of Al and SiC. Niu et al. (2006) reported that the welding of Al-SiC can be successfully realized by using laser welding. It is possible to control the geometry of weld bead by precisely controlling the laser output parameters (Yue et al., 1996). However, the reaction between SiC and Al in the weld zone generates a formation of needle-like brittle Al4C3, which can degrade the mechanical properties of Al-SiC composite. It dissolves in aqueous environments resulting in a loss of integrity to the weld zone (Lienert et al., 1993). In addition, the Al4C3 formation reduces the toughness of the weld bead, even it results in the increase of hardness (Bassani et al., 2007). Therefore, preventing the formation of the Al4C3 during welding process is important for successful welding of Al-SiC composites (Wang et al., 2000).

There are severe heterogeneity of the material structure and the great difference in both physical and chemical properties between the aluminum matrix and graphite particles. This point would make the welding of STC Al-Gr composites to a challenging technological problem for applying the laser welding technique in the industrial application. In this study, the pulsed Nd:YAG laser was used as a laser source. Square shape pulses are the standard output of this laser with the constant power distribution during its duration time. By utilizing the ability of pulsed Nd:YAG laser to shape the temporal power profile of each pulse offers the high flexibility in optimising the weld parameters. This pulse waveform allows the control of penetration, melt pool geometry and keyhole formation. It also has been reported that pulse waveform to be an effective method to reduce or eliminate the weld defects (Zhang et al., 2008).

#### **3. Experimental procedures**

164 Nd YAG Laser

allow melting and solidification consecutively. The great advantage of laser welding is keyhole effect by very high power densities involved in the laser welding, which makes

Material properties Value

zz 8×10¯6/K

Thermal conductivity kxx , kyy 450 W/(m·K) kzz 40 W/(m·K) Specific heat c 810 J/(kg·K) Density ρ 2450 kg/m³ Thermal expansion coefficient xx , yy 17×10¯6/K

Bending strength Mxx 65 MPa Mzz 7 MPa Tensile strength σ 49 MPa Young modulus G 31 GPa Electrical resistivity R 17×10¯8 Ωm

Fig. 3. Schematic illustration and optical micrographs of STC Al-Gr composite

Since the STC Al-Gr composite is a new composite in MMCs, no work has been reported concerning the welding behaviour on this composite by laser welding technique. There are several published papers related to the laser welding of MMCs, especially in the composites with combination of Al and SiC. Niu et al. (2006) reported that the welding of Al-SiC can be successfully realized by using laser welding. It is possible to control the geometry of weld bead by precisely controlling the laser output parameters (Yue et al., 1996). However, the reaction between SiC and Al in the weld zone generates a formation of needle-like brittle Al4C3, which can degrade the mechanical properties of Al-SiC composite. It dissolves in aqueous environments resulting in a loss of integrity to the weld zone (Lienert et al., 1993).

laser welding technique widely used in many industrial productions.

Table 1. Material properties of STC Al-Gr composite

A schematic diagram of experimental setup is shown in Figure 4. In this study, a pulsed Nd:YAG laser (LASAG SLS200 CL8) of 1064 nm in wavelength was used as a laser source. The laser beam was collimated to 15 mm in diameter and delivered by an optical fiber of 50 μm core diameter. The collimator was installed between the optical fiber and the bending mirror, and the collimated laser beam was focused on the specimen surface by a lens of 50 mm in focal length. In order to avoid the back-reflection of incident laser beam, the processing head was aligned by 10 degrees to the perpendicular axis of the specimen surface. The irradiation experiments were carried out in two sections, namely the bead-onplate welding of STC Al-Gr composite, and the overlap welding of pure aluminum and STC Al-Gr composite. The welding experiments were done in a shielding gas of nitrogen with 13L/min flow rate. The stage controller could determine the movement of X-Y-Z stage, and it also synchronized the laser pulse.

Fig. 4. Schematic diagram of experimental setup

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 167

specimens was designed in the analysis model in order to reduce the calculation time. In addition, the half model with symmetric conditions is used in the analysis. As the heat source of laser beam is symmetric in the x-z plane, only half the heat source is considered. The heat source comprises a Gaussian plane heat source on the top surface and a conical shape heat source along the thickness of the specimen. To simplify the analysis, it was assumed that the alignment of laser spot was perpendicular to the specimen surface and a single spot of laser irradiation was utilized in the heat conduction analysis. The convective heat transfer condition of air was considered after the set time of laser irradiation. Except for a laser beam irradiated area, the convective heat transfer condition of air was also assumed.

The main FEM analytical conditions are shown in Table 2.

Fig. 6. Finite element models for investigation of temperature distributions

Parameter/condition Bead-on-plate

Beam diameter d 50 μm Heat transfer coefficient h 10 W/(m²·K) Room temperature θroom 293 K

Table 2. FEM analysis conditions

**5. Bead-on-plate welding**

Laser power P 20, 30, 40 W 400 W Pulse width τ 1.5 ms 2.7, 3.0, 5.0 ms

In the bead-on-plate welding experiment, the STC Al-Gr composite of 1 mm thickness was used as the specimen. The specimen was mounted by clamping fixture with the aluminaceramic plates, which are located on the bottom surface of specimen to minimize the heat loss during welding experiments. The rectangular shape of laser pulse was fixed as the standard output of laser source. The experiments were conducted with laser power up to 50 W, pulse width from 1 to 5 ms and scanning velocities varied between 12 to 72 mm/min.

welding

Overlap welding

After the laser welding, the sectioned surface of welded specimen was ground, polished and etched for the observation of weld bead by an optical microscope, scanning electron microscope (SEM) and energy dispersive spectroscopy (EDS). In addition, the shearing test was carried out to measure the shear strength of the overlap welded joints with a Shimadzu EZ-L test machine. The specimen for shear strength of overlap welding was designed with five seam lines and 1.5 mm distance between each line as shown in Figure 5. The cross-head speed was set to 0.5 mm/min. The specimen was gripped by the clampers, which are placed in the fixture blocks. Then, a shear load was slowly increased at the suitable increments by the mechanical lever system until the welded joint of specimen was fractured. The shear strength was calculated by using fracture load and welding area. The value of shear strength was the average of three specimens. Then, the fracture surfaces were examined with SEM and EDS analysis.

Fig. 5. Configuration of shearing test specimen

#### **4. Thermal analysis by finite element method**

In this study, the further analysis of the welding phenomenon was discussed by the heat conduction analysis with finite element method (FEM) to simulate the thermal process of laser welding. The analysis model is based on the fundamental heat transfer for the laser welding process. Based on the first law of thermodynamics, the heat flow in a threedimensional solid can be expressed as equation (1) (Kannatey-Asibu, 2009)

$$\boldsymbol{\rho} \cdot \mathbf{c}(\boldsymbol{\theta}) \frac{\delta \boldsymbol{\theta}}{\delta \mathbf{t}} = \frac{\boldsymbol{\delta}}{\delta \mathbf{x}} \bigg[ \mathbf{k}(\boldsymbol{\theta}) \frac{\delta \boldsymbol{\theta}}{\delta \mathbf{x}} \bigg] + \frac{\boldsymbol{\delta}}{\delta \mathbf{y}} \bigg[ \mathbf{k}(\boldsymbol{\theta}) \frac{\delta \boldsymbol{\theta}}{\delta \mathbf{y}} \bigg] + \frac{\boldsymbol{\delta}}{\delta \mathbf{z}} \bigg[ \mathbf{k}(\boldsymbol{\theta}) \frac{\delta \boldsymbol{\theta}}{\delta \mathbf{z}} \bigg] + \mathbf{Q}\_{\mathbf{v}} \tag{1}$$

where ρ, c(θ), k(θ) and Qv are the material density, the temperature dependent specific heat, the temperature dependent thermal conductivity and the volumetric heat source term which varies with a laser power. θ=θ(x,y,z,t) is the resulting three-dimensional and temporal temperature distribution in the material. t is time, and x, y, z are the spatial Cartesian coordinates.

Figure 6 shows the developed finite element models for bead-on-plate welding and overlap welding. Fine mesh resolution is given at and near the heat source, while a fairly coarse mesh density is considered at the region far from the heat source. A portion of the

After the laser welding, the sectioned surface of welded specimen was ground, polished and etched for the observation of weld bead by an optical microscope, scanning electron microscope (SEM) and energy dispersive spectroscopy (EDS). In addition, the shearing test was carried out to measure the shear strength of the overlap welded joints with a Shimadzu EZ-L test machine. The specimen for shear strength of overlap welding was designed with five seam lines and 1.5 mm distance between each line as shown in Figure 5. The cross-head speed was set to 0.5 mm/min. The specimen was gripped by the clampers, which are placed in the fixture blocks. Then, a shear load was slowly increased at the suitable increments by the mechanical lever system until the welded joint of specimen was fractured. The shear strength was calculated by using fracture load and welding area. The value of shear strength was the average of three specimens. Then, the fracture surfaces were examined with SEM

In this study, the further analysis of the welding phenomenon was discussed by the heat conduction analysis with finite element method (FEM) to simulate the thermal process of laser welding. The analysis model is based on the fundamental heat transfer for the laser welding process. Based on the first law of thermodynamics, the heat flow in a three-

> <sup>v</sup> δθ δ δθ δ δθ δ δθ <sup>ρ</sup> <sup>c</sup> <sup>θ</sup> <sup>k</sup> <sup>θ</sup> <sup>k</sup> <sup>θ</sup> <sup>k</sup> <sup>θ</sup> <sup>Q</sup> <sup>δ</sup><sup>t</sup> <sup>δ</sup><sup>x</sup> <sup>δ</sup><sup>x</sup> <sup>δ</sup><sup>y</sup> <sup>δ</sup><sup>y</sup> <sup>δ</sup><sup>z</sup> <sup>δ</sup><sup>z</sup>

where ρ, c(θ), k(θ) and Qv are the material density, the temperature dependent specific heat, the temperature dependent thermal conductivity and the volumetric heat source term which varies with a laser power. θ=θ(x,y,z,t) is the resulting three-dimensional and temporal temperature distribution in the material. t is time, and x, y, z are the spatial Cartesian

Figure 6 shows the developed finite element models for bead-on-plate welding and overlap welding. Fine mesh resolution is given at and near the heat source, while a fairly coarse mesh density is considered at the region far from the heat source. A portion of the

(1)

dimensional solid can be expressed as equation (1) (Kannatey-Asibu, 2009)

and EDS analysis.

coordinates.

Fig. 5. Configuration of shearing test specimen

**4. Thermal analysis by finite element method**

specimens was designed in the analysis model in order to reduce the calculation time. In addition, the half model with symmetric conditions is used in the analysis. As the heat source of laser beam is symmetric in the x-z plane, only half the heat source is considered. The heat source comprises a Gaussian plane heat source on the top surface and a conical shape heat source along the thickness of the specimen. To simplify the analysis, it was assumed that the alignment of laser spot was perpendicular to the specimen surface and a single spot of laser irradiation was utilized in the heat conduction analysis. The convective heat transfer condition of air was considered after the set time of laser irradiation. Except for a laser beam irradiated area, the convective heat transfer condition of air was also assumed. The main FEM analytical conditions are shown in Table 2.

Fig. 6. Finite element models for investigation of temperature distributions

#### **5. Bead-on-plate welding**

In the bead-on-plate welding experiment, the STC Al-Gr composite of 1 mm thickness was used as the specimen. The specimen was mounted by clamping fixture with the aluminaceramic plates, which are located on the bottom surface of specimen to minimize the heat loss during welding experiments. The rectangular shape of laser pulse was fixed as the standard output of laser source. The experiments were conducted with laser power up to 50 W, pulse width from 1 to 5 ms and scanning velocities varied between 12 to 72 mm/min.

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 169

2002) compared with the melting point of aluminum (θmelt: 933 K). In addition, since the graphite is soft and brittle material, the graphite would be smashed only by direct irradiated

In the second series of welding experiments, the influence of pulse width was investigated. Figure 9 shows the appearance of weld beads under the various pulse widths at a constant laser power. It shows that the longer pulse width would generate more molten volume because the weld pool remains for an extended period of time. The widest welds are generated by the longer pulse width, which is considered due to the longer heating time. As a result, the irregular shapes and large size of cavities or blowholes are generated periodically when the pulse width was more than 1 ms. The tendency of blowhole occurrence and weld bead imperfections were drastically increased in the case of longer pulse width due to instabilities of the keyhole. In addition, the adverse effect of long pulse width on the weld penetration was believed due to a higher proportion of heat conducted laterally into the specimen. The results reflect that the higher energy input would enlarge processed zone. As mentioned later in section 5.2, based on the thermal analysis by FEM, a very short pulse width is enough to initiate the melting of aluminum and the evaporation of

laser beam to form the small particles during keyhole welding process.

Fig. 8. SEM photographs and element mapping of Al and C

graphite on the welding of Al-Gr composite.

Fig. 9. Influence of pulse width on the top surface of weld bead

In order to keep the continuity of the penetration, the overlapping ratio was kept constant at 20 %. An increase in scanning velocity would increase the pulse repetition rate in the pulsed

#### **5.1 Influence of welding parameters**

In the first experiment, the welding condition on the top surface of Al-Gr composite was investigated by changing the power of laser pulse. The pulse width, pulse repetition rate and scanning velocity were kept constant at 1 ms, 5 Hz and 12 mm/min, respectively. The results of this experiment are shown in Figure 7. It can be noticed that the laser power significantly affected the welding condition. As shown in figure, when the laser power is insufficient high, the molten zone was discontinuous. At the lower 20 W of laser power, the molten zone can be observed only on the aluminum material, while the graphite material is non-molten phase. It is consider that the insufficient laser energy would not react the graphite material.

It is assumed that too high laser power leads to the generation of weld defects with increasing the power density. In practice, the power density below 107 W/cm2 is generally advisable to avoid severe ejection of molten material (Cao et al., 2003). At the 50 W of laser power it seems too high power density resulting blowholes and a poor quality of weld surface. The power density on the specimen surface can be reduced to prevent the porosities at the weld pool surface by reducing the laser power. Therefore, the acceptable condition of weld surface is obtained between 30 W to 40 W of laser power. However, since the graphite element only can be melted under the high pressure with high temperature (Kirillin & Kostanovskii, 2003), the existence of this element in the molten zone of Al-Gr composite is important to identify. It might be considered that the required laser power to generate molten zone in this composite is relatively smaller than the welding of pure aluminum, even it has a higher thermal conductivity.

Fig. 7. Influence of laser power on the top surface of weld bead

In order to identify the presence of graphite element in the molten zone, the energydispersive X-ray spectroscopy (EDS) technique was used for the elemental analysis. Figure 8 shows the distribution map of elements Al and C on the top surface of molten zone. As shown in the figure, the element Al is the main composition detected in the molten zone. However, the EDS mapping shows that only small particles of graphite were observed on the top of the solidified aluminum. Judging from the distribution maps of elements, it can be presumed that the graphite was not mixture with aluminum, and it was ruptured into small particles. Even the aluminum material of Al-Gr composite was melted with the low laser power, the graphite material could not be melted. Because, the graphite only can be melted at the high temperature (θmelt: 4765 K) under the high pressure (Fmelt: 100 bar) (Joseph et al.,

In the first experiment, the welding condition on the top surface of Al-Gr composite was investigated by changing the power of laser pulse. The pulse width, pulse repetition rate and scanning velocity were kept constant at 1 ms, 5 Hz and 12 mm/min, respectively. The results of this experiment are shown in Figure 7. It can be noticed that the laser power significantly affected the welding condition. As shown in figure, when the laser power is insufficient high, the molten zone was discontinuous. At the lower 20 W of laser power, the molten zone can be observed only on the aluminum material, while the graphite material is non-molten phase. It is consider that the insufficient laser energy would not react the

It is assumed that too high laser power leads to the generation of weld defects with increasing the power density. In practice, the power density below 107 W/cm2 is generally advisable to avoid severe ejection of molten material (Cao et al., 2003). At the 50 W of laser power it seems too high power density resulting blowholes and a poor quality of weld surface. The power density on the specimen surface can be reduced to prevent the porosities at the weld pool surface by reducing the laser power. Therefore, the acceptable condition of weld surface is obtained between 30 W to 40 W of laser power. However, since the graphite element only can be melted under the high pressure with high temperature (Kirillin & Kostanovskii, 2003), the existence of this element in the molten zone of Al-Gr composite is important to identify. It might be considered that the required laser power to generate molten zone in this composite is relatively smaller than the welding of pure aluminum, even

In order to identify the presence of graphite element in the molten zone, the energydispersive X-ray spectroscopy (EDS) technique was used for the elemental analysis. Figure 8 shows the distribution map of elements Al and C on the top surface of molten zone. As shown in the figure, the element Al is the main composition detected in the molten zone. However, the EDS mapping shows that only small particles of graphite were observed on the top of the solidified aluminum. Judging from the distribution maps of elements, it can be presumed that the graphite was not mixture with aluminum, and it was ruptured into small particles. Even the aluminum material of Al-Gr composite was melted with the low laser power, the graphite material could not be melted. Because, the graphite only can be melted at the high temperature (θmelt: 4765 K) under the high pressure (Fmelt: 100 bar) (Joseph et al.,

**5.1 Influence of welding parameters**

it has a higher thermal conductivity.

Fig. 7. Influence of laser power on the top surface of weld bead

graphite material.

2002) compared with the melting point of aluminum (θmelt: 933 K). In addition, since the graphite is soft and brittle material, the graphite would be smashed only by direct irradiated laser beam to form the small particles during keyhole welding process.

Fig. 8. SEM photographs and element mapping of Al and C

In the second series of welding experiments, the influence of pulse width was investigated. Figure 9 shows the appearance of weld beads under the various pulse widths at a constant laser power. It shows that the longer pulse width would generate more molten volume because the weld pool remains for an extended period of time. The widest welds are generated by the longer pulse width, which is considered due to the longer heating time. As a result, the irregular shapes and large size of cavities or blowholes are generated periodically when the pulse width was more than 1 ms. The tendency of blowhole occurrence and weld bead imperfections were drastically increased in the case of longer pulse width due to instabilities of the keyhole. In addition, the adverse effect of long pulse width on the weld penetration was believed due to a higher proportion of heat conducted laterally into the specimen. The results reflect that the higher energy input would enlarge processed zone. As mentioned later in section 5.2, based on the thermal analysis by FEM, a very short pulse width is enough to initiate the melting of aluminum and the evaporation of graphite on the welding of Al-Gr composite.

Fig. 9. Influence of pulse width on the top surface of weld bead

In order to keep the continuity of the penetration, the overlapping ratio was kept constant at 20 %. An increase in scanning velocity would increase the pulse repetition rate in the pulsed

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 171

Fig. 11. Spatial temperature distributions in bead-on-plate welding

Fig. 12. Temperature histories in bead-on-plate welding

were utilized as the output of laser source.

In the overlap welding experiment, the pure aluminum sheet of 0.3 mm thickness and the STC Al-Gr composite of 1 mm thickness were used as the specimens. In order to achieve an optimum welding condition, the rectangular and controlled pulse waveforms of laser pulses

**6. Overlap welding**

laser welding. The influence of scanning velocity on the top surface of weld bead was shown in Figure 10. It is noted that with the decrease of scanning velocity led to the enhancement of the molten zone surface, in which the weld seam surface showed an uniform surface ripple formation without blowholes. It can be seen that the effect of the remaining energy of the previous pulse is significant to interact the Al-Gr composite by the subsequent laser pulse. However, the size of molten zone was wider at the higher scanning velocity with the blowholes and poor quality of weld surface. Because the scanning velocity matches an opposite with the heat input, and the increasing of the scanning velocity means the decreasing of the average power per unit weld length exerted on the welding line, thereby producing a small amount of intermixed melt. Therefore, the lower scanning velocity is required to produce acceptable welded joints and maintain welding quality.

Fig. 10. Influence of scanning velocity on the top surface of weld bead

#### **5.2 Temperature field induced in STC Al-Gr composite**

A thermal analysis was carried out, in order to study the temperature field induced in the welding process of Al-Gr composite**.** Figure 11 and 12 show the influence of laser power and pulse width on temperature distribution of Al-Gr composite spatially and temporally. It can be clearly clarified that the absorbed energy on the Al-Gr composite was quickly removed in x-direction compared to the *z*-direction due to the higher thermal conductivity in x-direction (kxx: 450 W/(mK)), which is more than ten times of the thermal conductivity in z-direction (kzz: 40 W/(mK)). It can be seen that the increase of laser power and pulse width could enable an elevated temperature. This can be described that an enhanced laser energy input is absorbed by Al-Gr composite. In particular, when the low laser power of 20 W and 30 W are applied, the maximum temperatures are approximately 2400 K and 3500 K, respectively, which reaches to the melting and evaporation temperature of aluminum material below evaporation temeparture of graphite material. At the higher laser power of 40 W, the evaporation of graphite material occurs because the peak temperature of 4607 K exceeds the evaporation temperature of graphite. However, the influence of pulse width shows the slight increasing on the temperature after the melting of aluminum or evaporation of graphite was achieved. According to the thermal analysis results, it can be concluded that the laser power is significantly influenced on the evaporation of graphite. Judging from these results, a very short pulse width is required to initiate the melting of aluminum and the evaporation of graphite on the welding of Al-Gr composite.

Fig. 11. Spatial temperature distributions in bead-on-plate welding

Fig. 12. Temperature histories in bead-on-plate welding

#### **6. Overlap welding**

170 Nd YAG Laser

laser welding. The influence of scanning velocity on the top surface of weld bead was shown in Figure 10. It is noted that with the decrease of scanning velocity led to the enhancement of the molten zone surface, in which the weld seam surface showed an uniform surface ripple formation without blowholes. It can be seen that the effect of the remaining energy of the previous pulse is significant to interact the Al-Gr composite by the subsequent laser pulse. However, the size of molten zone was wider at the higher scanning velocity with the blowholes and poor quality of weld surface. Because the scanning velocity matches an opposite with the heat input, and the increasing of the scanning velocity means the decreasing of the average power per unit weld length exerted on the welding line, thereby producing a small amount of intermixed melt. Therefore, the lower scanning velocity is required to produce acceptable welded joints and maintain welding quality.

Fig. 10. Influence of scanning velocity on the top surface of weld bead

A thermal analysis was carried out, in order to study the temperature field induced in the welding process of Al-Gr composite**.** Figure 11 and 12 show the influence of laser power and pulse width on temperature distribution of Al-Gr composite spatially and temporally. It can be clearly clarified that the absorbed energy on the Al-Gr composite was quickly removed in x-direction compared to the *z*-direction due to the higher thermal conductivity in x-direction (kxx: 450 W/(mK)), which is more than ten times of the thermal conductivity in z-direction (kzz: 40 W/(mK)). It can be seen that the increase of laser power and pulse width could enable an elevated temperature. This can be described that an enhanced laser energy input is absorbed by Al-Gr composite. In particular, when the low laser power of 20 W and 30 W are applied, the maximum temperatures are approximately 2400 K and 3500 K, respectively, which reaches to the melting and evaporation temperature of aluminum material below evaporation temeparture of graphite material. At the higher laser power of 40 W, the evaporation of graphite material occurs because the peak temperature of 4607 K exceeds the evaporation temperature of graphite. However, the influence of pulse width shows the slight increasing on the temperature after the melting of aluminum or evaporation of graphite was achieved. According to the thermal analysis results, it can be concluded that the laser power is significantly influenced on the evaporation of graphite. Judging from these results, a very short pulse width is required to initiate the melting of aluminum and

**5.2 Temperature field induced in STC Al-Gr composite**

the evaporation of graphite on the welding of Al-Gr composite.

In the overlap welding experiment, the pure aluminum sheet of 0.3 mm thickness and the STC Al-Gr composite of 1 mm thickness were used as the specimens. In order to achieve an optimum welding condition, the rectangular and controlled pulse waveforms of laser pulses were utilized as the output of laser source.

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 173

joining part between pure aluminum plate and Al-Gr composite was obtained with porosity and bump defects. According to the observation results of cross-section in the welding with rectangular pulse waveform, the pulse width of 0.6 ms was better condition

As mentioned in the above observation results of normal rectangular pulse waveform, the bump and porosity defects were the major problem in the overlap welding between pure aluminum and Al-Gr composite. In order to overcome these problems, a controlled pulse waveform is discussed, since it is considered that an appropriate controlled laser pulse waveform could generate a better welded joint. Figure 14 shows the controlled pulse waveform named as a spike pulse waveform. The spike pulse waveform is divided into two phases. At the phase one, the laser power of 400 W and pulse width of 0.6 ms were selected according to the previous experimental results of rectangular pulse waveform. Phase two is a subsequent function of the phase one by adding the heat to melt the bump generated in the phase one. This re-melting process is intended to remove the porosity in order to obtain a better joining state. Therefore, the experiments were carried out to discuss an appropriate

value of laser power P2 and time period t2 at the phase two of spike pulse waveform.

Fig. 14. Schematic illustration of welding process with spike pulse waveform

The pulse width was set less than 5 ms due to the limitation of maximum pulse width in the laser system used in this study. Figure 16 shows the cross-section views for various pulse widths under the same laser power of phase two to define the appropriate value of pulse width at the phase two for the spike pulse waveform. An appropriate irradiation time is necessary to melt and reduce a bump during a welding process. Within the range 1.5-2.5 ms

From the cross-section observation as shows in Figure 15, it was seen that the porosity and bump have remained at a range of laser power 50-125 W, while an undercut was generated at the laser power more than 200W. The observation results of laser power 150 W and 175 W showed acceptable joining conditions, and 175 W was selected as the laser power of phase two (P2) for the deeper penetration and less porosity defect. The next experiment was carried out to define an optimum pulse width to set the time period t2 at the phase two of

with the deeper penetration, smaller porosities and bumps.

**6.2 Controlling pulse waveform**

spike pulse waveform as shown in Figure 14.

#### **6.1 Rectangular pulse waveform**

The variations of laser power from 300 to 550 W with constant 1 ms pulse width were irradiated on the specimen of 0.3 mm thickness aluminum and 1.0 mm Al-Gr composite plate as shown in Figure 13. It could be seen that the laser pulse energy less than 400 mJ/pulse was insufficient to form a molten geometry for joining both materials, in which the penetration depth was shallow. The higher pulse energy was crucial to achieve the sufficient penetration depth and control the formation of molten geometry. However, the cross-section view above 450 mJ/pulse showed signs of porosity and bump defects. Moreover, the undercut defect was observed at the higher pulse energy of 550 mJ/pulse. Judging from these observation results, it was confirmed that the higher laser power (more than 400 W) causes destructive effects, and lower laser power (less than 400 W) will restrict the penetration depth and joining. From these viewpoints mentioned above, the next welding experiments were carried out under the constant 400 W laser power with various pulse widths in the rectangular pulse waveform.

Fig. 13. Welding results with rectangular pulse waveform

Figure 13 shows welding results with the rectangular pulse waveform for various pulse widths and laser powers. As shown in the figure, the width of laser pulses had a significant role on the penetration depth. The weld joint between aluminum plate and Al-Gr composite can be seen at the pulse widths more than 0.3 ms. The penetration depth and bead width gradually increased with increasing the pulse width. However, these increments led to the increase of bump size in the weld joint. On the other hand, the penetration depth increased gradually with increasing the laser pulse energy, and a joining part between pure aluminum plate and Al-Gr composite was obtained with porosity and bump defects. According to the observation results of cross-section in the welding with rectangular pulse waveform, the pulse width of 0.6 ms was better condition with the deeper penetration, smaller porosities and bumps.

#### **6.2 Controlling pulse waveform**

172 Nd YAG Laser

The variations of laser power from 300 to 550 W with constant 1 ms pulse width were irradiated on the specimen of 0.3 mm thickness aluminum and 1.0 mm Al-Gr composite plate as shown in Figure 13. It could be seen that the laser pulse energy less than 400 mJ/pulse was insufficient to form a molten geometry for joining both materials, in which the penetration depth was shallow. The higher pulse energy was crucial to achieve the sufficient penetration depth and control the formation of molten geometry. However, the cross-section view above 450 mJ/pulse showed signs of porosity and bump defects. Moreover, the undercut defect was observed at the higher pulse energy of 550 mJ/pulse. Judging from these observation results, it was confirmed that the higher laser power (more than 400 W) causes destructive effects, and lower laser power (less than 400 W) will restrict the penetration depth and joining. From these viewpoints mentioned above, the next welding experiments were carried out under the constant 400 W laser power with various

**6.1 Rectangular pulse waveform**

pulse widths in the rectangular pulse waveform.

Fig. 13. Welding results with rectangular pulse waveform

Figure 13 shows welding results with the rectangular pulse waveform for various pulse widths and laser powers. As shown in the figure, the width of laser pulses had a significant role on the penetration depth. The weld joint between aluminum plate and Al-Gr composite can be seen at the pulse widths more than 0.3 ms. The penetration depth and bead width gradually increased with increasing the pulse width. However, these increments led to the increase of bump size in the weld joint. On the other hand, the penetration depth increased gradually with increasing the laser pulse energy, and a As mentioned in the above observation results of normal rectangular pulse waveform, the bump and porosity defects were the major problem in the overlap welding between pure aluminum and Al-Gr composite. In order to overcome these problems, a controlled pulse waveform is discussed, since it is considered that an appropriate controlled laser pulse waveform could generate a better welded joint. Figure 14 shows the controlled pulse waveform named as a spike pulse waveform. The spike pulse waveform is divided into two phases. At the phase one, the laser power of 400 W and pulse width of 0.6 ms were selected according to the previous experimental results of rectangular pulse waveform. Phase two is a subsequent function of the phase one by adding the heat to melt the bump generated in the phase one. This re-melting process is intended to remove the porosity in order to obtain a better joining state. Therefore, the experiments were carried out to discuss an appropriate value of laser power P2 and time period t2 at the phase two of spike pulse waveform.

From the cross-section observation as shows in Figure 15, it was seen that the porosity and bump have remained at a range of laser power 50-125 W, while an undercut was generated at the laser power more than 200W. The observation results of laser power 150 W and 175 W showed acceptable joining conditions, and 175 W was selected as the laser power of phase two (P2) for the deeper penetration and less porosity defect. The next experiment was carried out to define an optimum pulse width to set the time period t2 at the phase two of spike pulse waveform as shown in Figure 14.

Fig. 14. Schematic illustration of welding process with spike pulse waveform

The pulse width was set less than 5 ms due to the limitation of maximum pulse width in the laser system used in this study. Figure 16 shows the cross-section views for various pulse widths under the same laser power of phase two to define the appropriate value of pulse width at the phase two for the spike pulse waveform. An appropriate irradiation time is necessary to melt and reduce a bump during a welding process. Within the range 1.5-2.5 ms

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 175

and Al-Gr composite. It is considered that the severe convection of graphite particles with aluminum materials might be occurred in the molten zone. In other words, when Al-Gr composite is evaporated, an over-pressure is developed in the keyhole. This phenomenon would cause graphite particles pushed up from the prior zone in Al-Gr composite to the zone in the pure aluminum with melting phase, and finally the carbon particles would be redistributed during the re-solidification. However, it could be seen that a lack of fusion defect was generated by the spike pulse waveform. It could be detected by using an accurate microscopic observation in the high magnification condition. The lack of fusion defect in the molten zone can be identified by its string-like appearance, and it had randomly oriented curvature. It is considered that the lack of fusion defect is not pore, since the carbon element

Fig. 17. SEM images and EDS mappings with spike pulse waveform

From the observation results by the spike pulse waveform, the use of controlled pulse waveform has a positive effect on molten zone to remove the porosity and minimize the bump. However, the rapid cooling after welding process would generate the coarse carbon particles and lack of fusion defect in the molten zone. Therefore, an improvement of heat input at the end of laser pulse is necessary by introducing an approach of rampdown on the phase two to relieve the internal stress during re-solidification process, which would generate a better welded joint. The new controlled pulse waveforms are expressed as the annealing pulse waveform, and the trailing pulse waveform is also shown in Figure 18, in which the setting profile and the actual signal of three controlled pulse waveforms are also shown. Since the phase two of spike pulse waveform generated a stable weld bead compared with the rectangular pulse waveform, the amount of energy at the phase two for the annealing and the trailing pulse waveforms was conducted under the same energy of 420 mJ. The main difference between the annealing and the trailing pulse waveforms are the laser power and irradiation time during the phase two. The annealing pulse waveform has the higher laser power with the shorter

could be seen in this area.

of pulse width (t2), the results show that the welding defect of bump was not removed. Moreover, it could be cleared that the longer irradiation time (t2: 3.5-5.0 ms) generated the larger size of bump. On the other hand, the pulse width t2 of 3 ms showed an acceptable penetration depth. Judging from these results, 3 ms is selected as an appropriate value of t2 at the phase two in the case of spike pulse waveform due to the stable penetration and smaller bump without porosities.

Fig. 15. Influence of laser power (P2) in phase 2 of spike pulse waveform

Fig. 16. Influence of pulse width (t2) in phase 2 of spike pulse waveform

In addition, the elemental analysis was performed to observe the element distribution at the welded parts using a SEM equipped with the energy dispersive spectroscopy (EDS). Figure 17 showed the results of EDS mapping analysis in the case of spike pulse waveform. The shape of molten zone at the Al-Gr composite can not be seen clearly because of too much carbon particles on the molten zone, which can not be melted during the welding process. It could be seen that the coarse carbon located at the welded joint between pure aluminum

of pulse width (t2), the results show that the welding defect of bump was not removed. Moreover, it could be cleared that the longer irradiation time (t2: 3.5-5.0 ms) generated the larger size of bump. On the other hand, the pulse width t2 of 3 ms showed an acceptable penetration depth. Judging from these results, 3 ms is selected as an appropriate value of t2 at the phase two in the case of spike pulse waveform due to the stable penetration and

Fig. 15. Influence of laser power (P2) in phase 2 of spike pulse waveform

Fig. 16. Influence of pulse width (t2) in phase 2 of spike pulse waveform

In addition, the elemental analysis was performed to observe the element distribution at the welded parts using a SEM equipped with the energy dispersive spectroscopy (EDS). Figure 17 showed the results of EDS mapping analysis in the case of spike pulse waveform. The shape of molten zone at the Al-Gr composite can not be seen clearly because of too much carbon particles on the molten zone, which can not be melted during the welding process. It could be seen that the coarse carbon located at the welded joint between pure aluminum

smaller bump without porosities.

and Al-Gr composite. It is considered that the severe convection of graphite particles with aluminum materials might be occurred in the molten zone. In other words, when Al-Gr composite is evaporated, an over-pressure is developed in the keyhole. This phenomenon would cause graphite particles pushed up from the prior zone in Al-Gr composite to the zone in the pure aluminum with melting phase, and finally the carbon particles would be redistributed during the re-solidification. However, it could be seen that a lack of fusion defect was generated by the spike pulse waveform. It could be detected by using an accurate microscopic observation in the high magnification condition. The lack of fusion defect in the molten zone can be identified by its string-like appearance, and it had randomly oriented curvature. It is considered that the lack of fusion defect is not pore, since the carbon element could be seen in this area.

Fig. 17. SEM images and EDS mappings with spike pulse waveform

From the observation results by the spike pulse waveform, the use of controlled pulse waveform has a positive effect on molten zone to remove the porosity and minimize the bump. However, the rapid cooling after welding process would generate the coarse carbon particles and lack of fusion defect in the molten zone. Therefore, an improvement of heat input at the end of laser pulse is necessary by introducing an approach of rampdown on the phase two to relieve the internal stress during re-solidification process, which would generate a better welded joint. The new controlled pulse waveforms are expressed as the annealing pulse waveform, and the trailing pulse waveform is also shown in Figure 18, in which the setting profile and the actual signal of three controlled pulse waveforms are also shown. Since the phase two of spike pulse waveform generated a stable weld bead compared with the rectangular pulse waveform, the amount of energy at the phase two for the annealing and the trailing pulse waveforms was conducted under the same energy of 420 mJ. The main difference between the annealing and the trailing pulse waveforms are the laser power and irradiation time during the phase two. The annealing pulse waveform has the higher laser power with the shorter

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 177

other words, the slow cooling process during re-solidification is necessary to avoid the appearance of carbon particles and minimize the lack of fusion defect in the molten zone.

Fig. 19. Temperature histories by three controlled pulse waveforms

Fig. 20. Spatial temperature distribution by three controlled pulse waveforms

interaction time, while the trailing pulse waveform has the lower laser power with the longer interaction time.

Fig. 18. Setting shape and actual signal of controlled pulse waveform

The difference of temperature change by spike, annealing and trailing pulse waveforms was investigated with the thermal calculation. Spike, annealing and trailing pulse waveforms are finished at the time of 3.0 ms, 2.7 ms and 5.0 ms, respectively. The temperature distributions were similar until 0.6 ms for three pulse waveforms, since they have similar pulse shape (laser power and pulse width) at the phase one. The main difference appeared at the pulse shape of phase two even under the same energy (Ep2: 420 mJ). Figure 19 and 20 show the temperature histories and distributions by these three controlled pulse waveforms. It can be seen that spike pulse waveform shows the constant temperature distribution during the phase two. Therefore, it is considered that the rapid cooling at 3 ms during the solidification generates internal stress, and it might cause the lack of fusion defect. From this disadvantage of spike pulse waveform, the slow cooling process is required during the solidification phase. As shown in the figures, the annealing pulse waveform shows the rapid decreasing of temperature distribution at the phase two. On the other hand, the trailing pulse waveform indicates the gradual decreasing of temperature distribution at the phase two, which means that slow cooling could be realized. In other words, the longer time of laser irradiation is useful at the phase two in order to overcome the lack of fusion problem.

SEM images and EDS mappings of cross-section with annealing and trailing pulse waveforms are shown in Figure 21. In the SEM image, the shape of molten zone can be seen clearly if there are existences of aluminum element on the Al-Gr composite material. Compared with the spike pulse waveform, the size of carbon particles was much smaller in the case of annealing pulse waveform. However, it shows that the lack of fusion was appeared around the carbon particles. In the case of trailing pulse waveform, it can be seen that the specimens have been molten well with an acceptable weld bead state without lack of fusion defect. It also shows that a better weld bead state was obtained by applying the trailing pulse waveform. Furthermore, the sufficient long time during welding process. In

interaction time, while the trailing pulse waveform has the lower laser power with the

The difference of temperature change by spike, annealing and trailing pulse waveforms was investigated with the thermal calculation. Spike, annealing and trailing pulse waveforms are finished at the time of 3.0 ms, 2.7 ms and 5.0 ms, respectively. The temperature distributions were similar until 0.6 ms for three pulse waveforms, since they have similar pulse shape (laser power and pulse width) at the phase one. The main difference appeared at the pulse shape of phase two even under the same energy (Ep2: 420 mJ). Figure 19 and 20 show the temperature histories and distributions by these three controlled pulse waveforms. It can be seen that spike pulse waveform shows the constant temperature distribution during the phase two. Therefore, it is considered that the rapid cooling at 3 ms during the solidification generates internal stress, and it might cause the lack of fusion defect. From this disadvantage of spike pulse waveform, the slow cooling process is required during the solidification phase. As shown in the figures, the annealing pulse waveform shows the rapid decreasing of temperature distribution at the phase two. On the other hand, the trailing pulse waveform indicates the gradual decreasing of temperature distribution at the phase two, which means that slow cooling could be realized. In other words, the longer time of laser irradiation is useful at the phase two in order to overcome the lack of

SEM images and EDS mappings of cross-section with annealing and trailing pulse waveforms are shown in Figure 21. In the SEM image, the shape of molten zone can be seen clearly if there are existences of aluminum element on the Al-Gr composite material. Compared with the spike pulse waveform, the size of carbon particles was much smaller in the case of annealing pulse waveform. However, it shows that the lack of fusion was appeared around the carbon particles. In the case of trailing pulse waveform, it can be seen that the specimens have been molten well with an acceptable weld bead state without lack of fusion defect. It also shows that a better weld bead state was obtained by applying the trailing pulse waveform. Furthermore, the sufficient long time during welding process. In

Fig. 18. Setting shape and actual signal of controlled pulse waveform

longer interaction time.

fusion problem.

other words, the slow cooling process during re-solidification is necessary to avoid the appearance of carbon particles and minimize the lack of fusion defect in the molten zone.

Fig. 19. Temperature histories by three controlled pulse waveforms

Fig. 20. Spatial temperature distribution by three controlled pulse waveforms

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 179

Fig. 22. Shear strength of the welding with and without the control of pulse waveforms (R:

Figure 24 shows the side fracture surface of aluminum and STC Al-Gr composite, which were obtained from the shearing test under the welding condition with trailing pulse waveform. It can be seen that the fracture appearances show the brittle fracture. This fracture along the welding interface would deteriorate the weld strength of aluminum and STC Al-Gr composite joint. Figure 24 also shows the distribution map of elements Al and C on the fracture zone, where spot analysis of points 1 to 10 are listed in Table 3. It can not be detected the aluminum carbide Al4C3 at the fracture zone, since the solid graphite can not be solidified into liquid aluminum. EDS mapping on the bottom surface of aluminum confirms that after the shearing test, the fine and coarse graphite particles were found sticks on

In the case of fracture zone on the top surface of STC Al-Gr composite, it also can be seen that the solidified aluminum was squeezed out towards the edge of weld joint region, which showed the aluminum and graphite can not mixed together between both materials.

Fig. 23. SEM photographs of fracture on the top of STC Al-Gr composite

melted aluminum without mixture with solidified aluminum.

Rectangular, S: Spike, A: Annealing, T: Trailing)

Fig. 21. SEM images and EDS mappings of cross-section with (a) annealing and (b) trailing pulse waveforms

#### **6.3 Evaluation of mechanical strength**

In order to evaluate the weld strength on the overlap welding with and without the control of pulse waveform under the constant laser pulse energy (Ep: 660 mJ/pulse), the shearing test was carried out. Figure 22 shows the shear strength for various laser pulse waveforms. As shown in the figure, the weld joint with controlled trailing pulse waveform indicated the greater weld strength compared to the uncontrolled pulse waveform. The lower strength of the uncontrolled pulse waveform is attributable to the fact that the existence of porosity reduced the strength of weld joint. Meanwhile, it shows that the spike pulse waveform was less significant to increase the weld strength compared to the uncontrolled pulse waveform. It can be noted that the lack of fusion defect was affected on the lower weld strength. Therefore, it is cleared that the the appropriate controlled laser pulse configurations are effective to improve the weld joint between aluminum and STC Al-Gr composite.

Figure 23 shows the fracture part on the top surface of Al-Gr composite after the shearing test. As can be seen from the figure, the fracture with uncontrolled pulse waveform occurred at the weld bead boundary in the Al-Gr composite and the presence of groove could be observed. It could be noticed that the existence of groove defect clearly influenced the weakness of weld joint strength. However, the fracture in the welding with the control of pulse waveform appeared inside the weld bead, which is located at the interface between pure aluminum and Al-Gr composite without the groove defect. In other words, the interfaces of aluminum and STC Al-Gr composite were expected to be the weak points of the weld joint. It is also considered that a weld joint without weld defects would increase the weld strength.

Fig. 21. SEM images and EDS mappings of cross-section with (a) annealing and (b) trailing

In order to evaluate the weld strength on the overlap welding with and without the control of pulse waveform under the constant laser pulse energy (Ep: 660 mJ/pulse), the shearing test was carried out. Figure 22 shows the shear strength for various laser pulse waveforms. As shown in the figure, the weld joint with controlled trailing pulse waveform indicated the greater weld strength compared to the uncontrolled pulse waveform. The lower strength of the uncontrolled pulse waveform is attributable to the fact that the existence of porosity reduced the strength of weld joint. Meanwhile, it shows that the spike pulse waveform was less significant to increase the weld strength compared to the uncontrolled pulse waveform. It can be noted that the lack of fusion defect was affected on the lower weld strength. Therefore, it is cleared that the the appropriate controlled laser pulse configurations are

effective to improve the weld joint between aluminum and STC Al-Gr composite.

Figure 23 shows the fracture part on the top surface of Al-Gr composite after the shearing test. As can be seen from the figure, the fracture with uncontrolled pulse waveform occurred at the weld bead boundary in the Al-Gr composite and the presence of groove could be observed. It could be noticed that the existence of groove defect clearly influenced the weakness of weld joint strength. However, the fracture in the welding with the control of pulse waveform appeared inside the weld bead, which is located at the interface between pure aluminum and Al-Gr composite without the groove defect. In other words, the interfaces of aluminum and STC Al-Gr composite were expected to be the weak points of the weld joint. It is also considered that a weld joint without weld

pulse waveforms

**6.3 Evaluation of mechanical strength**

defects would increase the weld strength.

Fig. 22. Shear strength of the welding with and without the control of pulse waveforms (R: Rectangular, S: Spike, A: Annealing, T: Trailing)

Fig. 23. SEM photographs of fracture on the top of STC Al-Gr composite

Figure 24 shows the side fracture surface of aluminum and STC Al-Gr composite, which were obtained from the shearing test under the welding condition with trailing pulse waveform. It can be seen that the fracture appearances show the brittle fracture. This fracture along the welding interface would deteriorate the weld strength of aluminum and STC Al-Gr composite joint. Figure 24 also shows the distribution map of elements Al and C on the fracture zone, where spot analysis of points 1 to 10 are listed in Table 3. It can not be detected the aluminum carbide Al4C3 at the fracture zone, since the solid graphite can not be solidified into liquid aluminum. EDS mapping on the bottom surface of aluminum confirms that after the shearing test, the fine and coarse graphite particles were found sticks on melted aluminum without mixture with solidified aluminum.

In the case of fracture zone on the top surface of STC Al-Gr composite, it also can be seen that the solidified aluminum was squeezed out towards the edge of weld joint region, which showed the aluminum and graphite can not mixed together between both materials.

Micro-Welding of Super Thermal Conductive Composite by Pulsed Nd:YAG Laser 181

the STC Al-Gr composite without evaporation of the graphite element in the composite. The graphite was not mixed with aluminum during welding process to prevent the formation of aluminum carbide, which can degrade the weld joint. The overlap welding of aluminum and STC Al-Gr composite was successfully carried out using an appropriate controlled pulse waveform. Porosity and bump were observed as remarkable weld defects in overlap welding without a control of laser pulse. The proper control of laser power and pulse width could perform a positive result with largely free of weld defects and a relatively small bump. The controlled pulse waveform with slow cooling at the end of laser pulse was essential to relieve internal stress during solidification, since the lack of fusion was observed on the joining zone due to the rapid cooling. The higher shearing strength could be obtained by the control of pulse waveform compared with the uncontrolled rectangular pulse

The authors are grateful to Shimane Institute for Industrial Technology for supplying super

Baraker, N.; Tzamtzis, S.; Dhindaw, B.K.; Patel, J.; Babu, N.H. and Fan, Z. (2009). Processing

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waveform.

**8. Acknowledgment**

**9. References**

thermal conductive aluminum-graphite composites.

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Furthermore, the formation of the Al4C3 during welding process was successfully prevented during the welding process, which could deteriorate the strength of weld joint. In addition, the SEM observation revealed the crack was propagated in the fracture zone, which restricts the further strength of weld joint. However, compared to the bending strength of STC Al-Gr composite which is 7 MPa in the thickness direction, the weld strength between the aluminum and STC Al-Gr composite showed the relatively higher strength. Therefore, it is clearly performed that the controlled laser pulse configurations are effective to produce a higher strength of joint for welding between an aluminum and STC Al-Gr composite.

Fig. 24. SEM photographs and EDS mappings of fracture with trailing pulse waveform on the (a) bottom surface of aluminum and (b) top surface of STC Al-Gr composite

Table 3. Element composition of points 1-10 in Figure 24

#### **7. Conclusion**

The pulsed Nd:YAG laser micro-welding of a super thermal conductive (STC) aluminumgraphite composite was experimentally and numerically investigated. In the bead-on-plate welding of STC Al-Gr composite, the laser power and pulse width had a great influence on the top surface condition of weld bead. Laser power more than 30 W was required to melt the STC Al-Gr composite without evaporation of the graphite element in the composite. The graphite was not mixed with aluminum during welding process to prevent the formation of aluminum carbide, which can degrade the weld joint. The overlap welding of aluminum and STC Al-Gr composite was successfully carried out using an appropriate controlled pulse waveform. Porosity and bump were observed as remarkable weld defects in overlap welding without a control of laser pulse. The proper control of laser power and pulse width could perform a positive result with largely free of weld defects and a relatively small bump. The controlled pulse waveform with slow cooling at the end of laser pulse was essential to relieve internal stress during solidification, since the lack of fusion was observed on the joining zone due to the rapid cooling. The higher shearing strength could be obtained by the control of pulse waveform compared with the uncontrolled rectangular pulse waveform.

#### **8. Acknowledgment**

The authors are grateful to Shimane Institute for Industrial Technology for supplying super thermal conductive aluminum-graphite composites.

#### **9. References**

180 Nd YAG Laser

Furthermore, the formation of the Al4C3 during welding process was successfully prevented during the welding process, which could deteriorate the strength of weld joint. In addition, the SEM observation revealed the crack was propagated in the fracture zone, which restricts the further strength of weld joint. However, compared to the bending strength of STC Al-Gr composite which is 7 MPa in the thickness direction, the weld strength between the aluminum and STC Al-Gr composite showed the relatively higher strength. Therefore, it is clearly performed that the controlled laser pulse configurations are effective to produce a

higher strength of joint for welding between an aluminum and STC Al-Gr composite.

Fig. 24. SEM photographs and EDS mappings of fracture with trailing pulse waveform on

Point No. 1 2 3 4 5 6 7 8 9 10 Al wt. % 49.67 30.94 62.84 41.16 64.45 52.45 35.68 48.66 64.65 48.01 C wt. % 50.33 69.06 37.16 58.84 35.55 47.55 64.32 51.34 35.35 51.99

The pulsed Nd:YAG laser micro-welding of a super thermal conductive (STC) aluminumgraphite composite was experimentally and numerically investigated. In the bead-on-plate welding of STC Al-Gr composite, the laser power and pulse width had a great influence on the top surface condition of weld bead. Laser power more than 30 W was required to melt

the (a) bottom surface of aluminum and (b) top surface of STC Al-Gr composite

Table 3. Element composition of points 1-10 in Figure 24

**7. Conclusion**

**0**

**10**

*Czech Republic*

**Highly Doped Nd:YAG Laser in Bounce Geometry**

Nanosecond and picosecond laser pulses play increasing role in a variety of applications, such as microsurgery, micromachining, ranging, optical parametric amplifiers pumping, remote sensing, nonlinear optics, etc. Efficient diode pumped solid state lasers represent widely used sources of such pulses. For applications such as laser ranging and some nonlinear optical effects investigation, the extremely short pulses are not needed and pulses in the range from tens of picoseconds to units of nanoseconds at repetition rates below 1 kHz are optimal.

In the diode pumped solid state laser technique several resonator configurations and pumping schemes are utilized. One of interesting setups with unique properties is bounce geometry resonator configuration (Alcock, 1997) when a slab active material is used and the laser beam experiences total internal reflection and amplification on the pumped crystal side. This arrangement has several advantages. The slab crystal shape allows good heat dissipation, because it is relatively thin and can be cooled from one or both sides. As a pump source, high power laser diode single bar or bar stacks can be used without bulky additional optics or radiation delivery fibers. In addition, laser resonator mode can be optimized to match the pumped region profile. Efficient operation of the bounce geometry laser pumped by the laser diode bar requires active medium with a high absorption coefficient for the pump radiation in order to obtain thin region with high gain near the pumping side. Therefore, materials as Nd:GdVO4 and Nd:YVO4 has challenged Nd:YAG with standard doping concentration of around 1 at.%. When recently Nd:YAG laser crystals with higher doping level (more than 2 at.%) became available (Urata, 2001), efficient operation of these crystals under bounce geometry has been demonstrated operating at 1.06 and 1.3 *μ*m (Sauder, 2006; 2007). Nd:YAG crystals have better thermal and mechanical properties, and better energy storage capacity due to the longer upper state life-time in comparison with vanadates giving potential to

Research and development of such laser systems is the subject of this chapter.

generate more energetic pulses in the Q-switched and mode-locked regime.

Highly-energetic pulses in the nanosecond range can be obtained from the oscillator using active or passive Q-switching, which is well-known and widespread technique, mostly used under continuous pumping. In order to obtain more energetic pulses, quasi-continuous (QCW) pumping can be employed. Using the 2 at.% Nd:YAG in bounce geometry under QCW pumping, passively Q-switched operation at 1.06 *μ*m has been reported with maximum

**1. Introduction**

**Under Quasi-Continuous Diode Pumping**

Michal Jelínek and Václav Kubeˇcek *Czech Technical University in Prague*

### **Highly Doped Nd:YAG Laser in Bounce Geometry Under Quasi-Continuous Diode Pumping**

Michal Jelínek and Václav Kubeˇcek *Czech Technical University in Prague Czech Republic*

#### **1. Introduction**

182 Nd YAG Laser

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Nanosecond and picosecond laser pulses play increasing role in a variety of applications, such as microsurgery, micromachining, ranging, optical parametric amplifiers pumping, remote sensing, nonlinear optics, etc. Efficient diode pumped solid state lasers represent widely used sources of such pulses. For applications such as laser ranging and some nonlinear optical effects investigation, the extremely short pulses are not needed and pulses in the range from tens of picoseconds to units of nanoseconds at repetition rates below 1 kHz are optimal. Research and development of such laser systems is the subject of this chapter.