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## **Meet the editor**

Dr. Wladislav Sudnik is a Professor of Welding at the Tula State University Welding Department. He received his Diploma in Welding Technology at Russian State Technological University, Moscow in 1965, his Doctorate in physics of metals at Tula State University in 1973, and his doctor habilitatus in Welding Technology at St. Petersburg State Polytechnic University in 1991.

From 1977-1978 he was a Researcher at the Max-Planck for Iron and Steel Research. From 1991 to 2001 joint researches were carried out on development of models and creation of the software for arc welding with ISF RWTH, Aachen and Daimler AG, Stuttgart. He is a first recipient of the Sossenheimer Software Innovation Award of International Institute of Welding in 2003 for his outstanding technical-scientific work concerning Modelling and Simulation in the field of welding and allied processes.

### Contents

#### **Preface** XI

#### **Part 1 Arc Welding Technology 1**

	- **Part 2 Arc Welding Automation 45**
	- **Part 3 Weldability of Metals and Alloys 107**

X Contents

Matthew Gonser and Theodore Hogan

### Preface

Ever since the invention of arc technology in 1870s and it's early use for welding lead during the manufacture of lead-acid batteries, advances in arc welding throughout the twentieth and twenty-first centuries have seen this form of processing applied to a range of industries and progress to become one of the most effective techniques in metals and alloys joining.

The objective of this book is to introduce relatively established methodologies and techniques which have been studied, developed and applied in industries or researches. State-of-the-art development aimed at improving technologies will be presented covering topics such as weldability, technology, automation, modelling, and measurement. This book also seeks to provide effective solutions to various applications for engineers and researchers who are interested in arc material processing.

This book is divided into 4 independent chapters corresponding to recent advances in this field.

The editor expresses thankfulness to all authors for the presented materials and their timely design, and also to the technical editor and to the book manager Mrs. Marija Radja - for the big work on preparation and the edition of this book.

Editor

**Prof. Dr. Wladislav Sudnik** R & E Center 'Computer Hi-Tech in Materials Joining' Welding Department Tula State University, Russian Federation

**Part 1**

**Arc Welding Technology**

### **Hardfacing by Plasma Transferred Arc Process**

Víctor Vergara Díaz1, Jair Carlos Dutra2 and Ana Sofia Climaco D'Oliveira3 *1University of Antofagasta, Mechanical Engineering Department 2University Federal de Santa Catarina, Mechanical Engineering Department 3University Federal do Paraná, Mechanical Engineering Department 1Chile 2,3Brasil*

#### **1. Introduction**

According to the literature, the plasma transferred arc welding process which employs the filler metal in wire form is known as Plasma Arc Welding (PAW) while that which employs powder filler material is generally referred to as Plasma Transferred Arc (PTA), Dai et al., 2008.

The PTA process can be considered a derivation of the PAW process. The similarities between the two processes can be observed in Figure 1. Both welding processes employ a non-consumable tungsten electrode located inside the torch, a water-cooled constrictor nozzle, shield gas for the protection of the molten pool, and the plasma gas. The difference between the two welding processes lies in the nature of the filler material, powder instead of wire, which requires a gas for its transport to the arc region. The diagram in Figure 1 shows the two processes with their differences and similarities.

The equipment required to carry out the deposition through the PTA plasma process is very similar to that used in PAW. When PAW is employed the equipment must be able to drive spooled wires of various gages and different materials, at constant or pulsed velocities. In the PTA plasma welding process, the filler material is used in the form of a powder, and specific powder feeding equipment is required to transport it to the voltaic arc to produce the coating. With respect to its application for coating, the PTA process is appropriate since it produces dilution values of the order of 6 to 10 % (Gatto, et al., 2009), much lower than those obtained with other arc soldering process which are around 20 to 25 %. The low distortion, the small zone affected by the heat and the refined microstructure are also features of this technique (Zhang, et al., 2008; Liu, et al., 2008).

In the PTA and PAW processes an inert gas is used as the plasma gas, which is forced to pass through the orifice of the constrictor nozzle, where the electrode is concentrically fixed. The shield gas passes through an external opening, concentric to the constrictor nozzle, effectively protecting the weld against contamination from atmospheric air (active or inert). On the other hand, in the PTA process a carrier gas is used to transport the filler material through flexible tubes to the constrictor nozzle, allowing its entrance into the plasma arc in a convergent form. The gas used for this purpose is generally argon.

PTA PROCESS PAW PROCESS

Fig. 1. Comparison of Plasma Transferred Arc processes PTA and PAW.

Given that the tungsten electrode lies within the constrictor nozzle of the welding torch, it is difficult to open the arc by contact, and thus equipment called a plasma module must be used to establish the arc opening. An electronic igniter provides voltage peaks between the tungsten electrode and constrictor nozzle, generating a small spark in this region. Thus, with the passage of the plasma gas a low intensity electric arc appears between the tungsten electrode and constrictor nozzle, called the pilot arc (non-transferred arc). The pilot arc forms a pathway of low electrical resistance between the tungsten electrode and the workpiece to be welded facilitating the establishment of the main arc when a power source is added.

In practice, the parameters which control the quality of the weld are the rate at which the material is added, the gas flow rate (shield gas, plasma gas, carrier gas), the weld current, the nozzle to workpiece distance (see below) and the welding speed.

The basic configuration of the constrictor nozzle is shown in Figure 2, where the parameters employed in the process are indicated. The distance from the external face of the constrictor nozzle to the substrate is called the nozzle to workpiece distance (NWD).

The recess (Rc) of the electrode is measured from the electrode tip to the external face of the constrictor nozzle. Alterations in the arc characteristics are influenced by this factor, which defines the degree of constriction and the rigidity of the plasma jet (Oliveira, 2001).

Oliveira (2001) studied the influence of the electrode recess of the plasma transferred arc process fed by wire in order to identify whether the degree of arc constriction influences the arc voltage. The results showed that, on average, a 2.4 V/mm variation in the voltage occurred as a function of the electrode recess.

Plasma gas flow

Electrode

PTA PROCESS PAW PROCESS

Given that the tungsten electrode lies within the constrictor nozzle of the welding torch, it is difficult to open the arc by contact, and thus equipment called a plasma module must be used to establish the arc opening. An electronic igniter provides voltage peaks between the tungsten electrode and constrictor nozzle, generating a small spark in this region. Thus, with the passage of the plasma gas a low intensity electric arc appears between the tungsten electrode and constrictor nozzle, called the pilot arc (non-transferred arc). The pilot arc forms a pathway of low electrical resistance between the tungsten electrode and the workpiece to be welded facilitating the establishment of the main arc when a power source

In practice, the parameters which control the quality of the weld are the rate at which the material is added, the gas flow rate (shield gas, plasma gas, carrier gas), the weld current,

The basic configuration of the constrictor nozzle is shown in Figure 2, where the parameters employed in the process are indicated. The distance from the external face of the constrictor

The recess (Rc) of the electrode is measured from the electrode tip to the external face of the constrictor nozzle. Alterations in the arc characteristics are influenced by this factor, which

Oliveira (2001) studied the influence of the electrode recess of the plasma transferred arc process fed by wire in order to identify whether the degree of arc constriction influences the arc voltage. The results showed that, on average, a 2.4 V/mm variation in the voltage

defines the degree of constriction and the rigidity of the plasma jet (Oliveira, 2001).

Fig. 1. Comparison of Plasma Transferred Arc processes PTA and PAW.

the nozzle to workpiece distance (see below) and the welding speed.

occurred as a function of the electrode recess.

nozzle to the substrate is called the nozzle to workpiece distance (NWD).

is added.

Substrate

Plasma arc

 Constrictor nozzle

Wire

Powder Shield gas

Fig. 2. Nozzle to workpiece distance (NWD) and electrode setback (Rc) (Vergara, 2005).

In general, the maximum and minimum values for the adjustment of the electrode recess vary according to the welding torch. The electrode recess of the welding torch PWM–300, manufactured by Thermal Dynamics Corporation, for instance, has a range of adjustment of 0.8 to 2.4 mm.

As the electrode recess is reduced, the weld bead width increases and weld beads with lower penetration depth are obtained. This variation in the geometric characteristics of the weld bead is due to a reduction in the constriction effect producing a larger area of incidence of the arc on the substrate.

The constrictor nozzle (made of copper), where the electrode is confined, has a central orifice through which the arc and all of the plasma gas volume pass. The diameter of the orifice of the constrictor nozzle has a great influence on the quality of the coating since this relationship is directly related to the width and penetration of the weld bead produced. An insufficient plasma gas flow rate affects the useful life of the constrictor nozzle since it leads to its wear. The weld current reduces as a function of the decrease in the diameter of the constricting orifice, due to an increase in the weld arc temperature.

The extent to which the nozzle to workpiece distance influences the coating is strongly dependent on the electrode recess in relation to the constrictor nozzle and the diameter of the constrictor orifice. The larger the electrode recess adopted and the smaller the constrictor orifice diameter the greater the effect of the arc constriction, making it more concentrated.

In the "melt–in" technique small electrode recess values are used, the arc being submitted to a low degree of collimation, assuming a conical form. In this situation, a variation in the nozzle to workpiece distance, even within normal limits, results in a change in the characteristics of the weld bead, in the same way as occurs in the GTAW process. Thus, the greater the nozzle to workpiece distance the lower the penetration and wider the width of the weld bead due to the increase in the area of incidence of the arc on the substrate.

Hallen et al. (1991) reported that to obtain a good deposition yield, the nozzle to workpiece distance should not be greater than 10 to 15 mm. At values higher than this range the efficiency of the shield gas is significantly reduced.

The authors of this paper have also reported results in relation to the nozzle to workpiece distance, for two values: 15 and 20 mm. The study showed that as the nozzle to workpiece distance increases the degree of dilution decreases.

The general objective of this study was to investigate the PAW and PTA welding processes with a view to their application in surface coating operations, particularly on hydraulic turbine blades worn by cavitation. This research was motivated by the observation that information is scare in relation to the benefits offered by the plasma welding process using powder instead of wire filler material in the application of coatings. The geometric characteristics of the weld beads, degree of dilution, hardness and microstructure were evaluated.

#### **2. Materials and Methods**

#### **2.1 Test bench**

Initially, a test bench was assembled based on equipment previously developed at LABSOLDA (Oliveira, 2001; Vergara, 2005) which allowed tests to be carried out on the plasma transferred arc welding process fed by wire. On the same test bench, a similar process fed by powder was assembled. The welding source was equipment which, via an interface, was connected to a PC. By way of a very versatile software program almost all of the process variables could be controlled.

Of the three gas circuits, that which received most attention was the plasma gas given its considerable relevance in terms of the quality of the deposits. A mass flow controller was used, in which the control is carried out electronically and the command signal is a reference voltage. The other gas flow circuits are simply monitored by electronic flow meters, however these are volumetric.

One of the fundamental parts of the equipment is the device known as the plasma module, which enables any version of plasma welding to be carried out based on conventional welding sources for GTAW or coated electrode. For the displacement of the welding torch an electronic device (Tartílope) was used. The system component which was integrally designed for this specific development was the powder feeding device, which functions through a combination of an endless screw and a gas flow as the powder carrying mechanisms. The weld torch was developed based on the plasma torch for keyhole welding. The great advantage of this lies in its multiprocess aspect which allows it to work with plasma employing powder or with conventional plasma. Also, the design adaptation allows the use of constrictor nozzles with different angles of convergence for the powder feeding. Initially, analysis was carried out on the torches to be used in this research. It was observed that the PTA torch had a nozzle with a constrictor diameter of 4.8 mm. In the case of the PAW torch, the manufacturer provides three nozzles with constrictor diameters of 2.4, 2.8 and 3.2 mm, which are designed according to the welding current to be applied.

In this case, the nozzle with the largest constrictor diameter available for the PAW torch was selected, that is, 3.2 mm.

Figure 3 shows a general view of the equipment developed, that which forms part of the test bench for the PAW and PTA welding processes being shown in the upper part of the figure.

Hallen et al. (1991) reported that to obtain a good deposition yield, the nozzle to workpiece distance should not be greater than 10 to 15 mm. At values higher than this range the

The authors of this paper have also reported results in relation to the nozzle to workpiece distance, for two values: 15 and 20 mm. The study showed that as the nozzle to workpiece

The general objective of this study was to investigate the PAW and PTA welding processes with a view to their application in surface coating operations, particularly on hydraulic turbine blades worn by cavitation. This research was motivated by the observation that information is scare in relation to the benefits offered by the plasma welding process using powder instead of wire filler material in the application of coatings. The geometric characteristics of the weld beads, degree of dilution, hardness and microstructure were

Initially, a test bench was assembled based on equipment previously developed at LABSOLDA (Oliveira, 2001; Vergara, 2005) which allowed tests to be carried out on the plasma transferred arc welding process fed by wire. On the same test bench, a similar process fed by powder was assembled. The welding source was equipment which, via an interface, was connected to a PC. By way of a very versatile software program almost all of

Of the three gas circuits, that which received most attention was the plasma gas given its considerable relevance in terms of the quality of the deposits. A mass flow controller was used, in which the control is carried out electronically and the command signal is a reference voltage. The other gas flow circuits are simply monitored by electronic flow meters,

One of the fundamental parts of the equipment is the device known as the plasma module, which enables any version of plasma welding to be carried out based on conventional welding sources for GTAW or coated electrode. For the displacement of the welding torch an electronic device (Tartílope) was used. The system component which was integrally designed for this specific development was the powder feeding device, which functions through a combination of an endless screw and a gas flow as the powder carrying mechanisms. The weld torch was developed based on the plasma torch for keyhole welding. The great advantage of this lies in its multiprocess aspect which allows it to work with plasma employing powder or with conventional plasma. Also, the design adaptation allows the use of constrictor nozzles with different angles of convergence for the powder feeding. Initially, analysis was carried out on the torches to be used in this research. It was observed that the PTA torch had a nozzle with a constrictor diameter of 4.8 mm. In the case of the PAW torch, the manufacturer provides three nozzles with constrictor diameters of 2.4, 2.8 and 3.2 mm, which are designed according to the welding

In this case, the nozzle with the largest constrictor diameter available for the PAW torch was

Figure 3 shows a general view of the equipment developed, that which forms part of the test bench for the PAW and PTA welding processes being shown in the upper part of the figure.

efficiency of the shield gas is significantly reduced.

distance increases the degree of dilution decreases.

evaluated.

**2.1 Test bench**

**2. Materials and Methods**

however these are volumetric.

current to be applied.

selected, that is, 3.2 mm.

the process variables could be controlled.

In this study argon with a purity of 99.99 % was used as the plasma, shield and carrier gases. A tungsten electrode with 2% thorium oxide (EWTh-2) and with a diameter of 4.8 mm was used. The angle of the electrode tip was maintained at 30º for all of the experiments.

Fig. 3. Test bench assembled at the welding laboratory. 1-Welding source; 2-Adapted plasma torch; 3-Plasma module; 4-Powder feeder; 5-Torch displacement system; 6-Digital gas meters; 7-Electronic gas valve; 8-Gases

#### **2.2 Constrictor nozzle in PTA process**

The configuration of the constrictor nozzle developed in this study included two conduits for the passage of the carrier gas, the role of which is to feed the powder to the plasma arc in a convergent form. Figure 4 shows a cross-section of the constrictor nozzle. At 60º the constrictor nozzle allows the entrance of powder directly into the molten pool, when a nozzle to workpiece distance of 10 mm is used.

Fig. 4. Cross-section of constrictor nozzle showing the entrance of the powder flow into the plasma arc. (Vergara, 2005).

#### **2.3 Characterization**

Deposits of the atomized alloy Stellite 6, Figure 5, were processed on carbon steel plates (class ABNT 1020; dimensions 12.5 x 60 x 155 mm), using a constant continuous current. Table 1 shows the chemical composition of the substrate. The chemical analysis of the different filler materials was carried out by optical emission spectrometry and the results are shown in Tables 2 and 3.

Single weld beads were deposited with the parameters indicated in Table 4 and samples were removed for their characterization. This table gives the operational parameters for the PTA and PAW plasma welding processes, in which there are parameters which could not remain constant in the two process, for example: nature of the filler material (in PAW wire and in PTA powder); wire speed (not required in PTA); carrier gas (not required in PAW); constrictor nozzle diameter (in PTA 4.8 mm and in PAW 3.2 mm).

Initially, the weld beads were submitted to visual inspection for the presence of welding defects, the degree of dilution was determined by the areas method using micrographs of the cross-sections of the deposits, etched with 6% nital. Profiles of the Vickers microhardness, with a load of 500g, enabled the evaluation of the uniformity of the weld beads processed, according to the procedure of the standard ABNT6672/81. The determination of the microhardness profiles, average of three measurements, was carried out at the center of the weld beads and in the region where they overlap. To determine the microstructure by optical microscopy a cross-section was prepared following standard procedures, the microstructure being revealed after electrolytic attack with oxalic acid.

Fig. 4. Cross-section of constrictor nozzle showing the entrance of the powder flow into the

Deposits of the atomized alloy Stellite 6, Figure 5, were processed on carbon steel plates (class ABNT 1020; dimensions 12.5 x 60 x 155 mm), using a constant continuous current. Table 1 shows the chemical composition of the substrate. The chemical analysis of the different filler materials was carried out by optical emission spectrometry and the results are

Single weld beads were deposited with the parameters indicated in Table 4 and samples were removed for their characterization. This table gives the operational parameters for the PTA and PAW plasma welding processes, in which there are parameters which could not remain constant in the two process, for example: nature of the filler material (in PAW wire and in PTA powder); wire speed (not required in PTA); carrier gas (not required in PAW);

Initially, the weld beads were submitted to visual inspection for the presence of welding defects, the degree of dilution was determined by the areas method using micrographs of the cross-sections of the deposits, etched with 6% nital. Profiles of the Vickers microhardness, with a load of 500g, enabled the evaluation of the uniformity of the weld beads processed, according to the procedure of the standard ABNT6672/81. The determination of the microhardness profiles, average of three measurements, was carried out at the center of the weld beads and in the region where they overlap. To determine the microstructure by optical microscopy a cross-section was prepared following standard procedures, the microstructure being revealed after electrolytic attack with oxalic acid.

constrictor nozzle diameter (in PTA 4.8 mm and in PAW 3.2 mm).

plasma arc. (Vergara, 2005).

**2.3 Characterization**

shown in Tables 2 and 3.

Fig. 5. Morphology of powder deposited by the PTA process (Stellite 6).

Thickness: 12.7 mm

Table 1. Chemical composition of the low carbon steel substrate.

Hardness: 38-47 Rc; Particle size: 45 to 150 µm; Density: 8.3 g/cm3

Table 2. Chemical composition of the filler material Stellite 6 in the form of a powder (BT-906)

Table 3. Chemical composition of filler material Stellite 6 in the form of steel (BT-906T).

Table 4. Welding variables and parameters.

#### **3. Results and discussion**

#### **3.1 General characteristics**

Figure 6 shows the external aspect of the beads where significant differences between them can be observed. The PTA process produced a better surface finish, better dilution, better wetting and wider width.

Figures 7 and 8 show cross-sections of the beads obtained using the two processes (PAW and PTA) where considerable differences in the penetration profile of the welds can be noted and Figure 9 shows the results for the geometric parameters of the beads, for the three levels of plasma gas flow rate tested in this study: 2.2; 2.4 and 3.0 l/min. On comparing the deposits obtained from the two processes it can be observed that the reinforcement and the penetration are always smaller in the PTA process (Figure 9). In the PTA process there was a significantly wider cord width, which is due to the use of a constrictor nozzle with a wider diameter.

The data shown in Figure 9 together with an analysis of the variance in Tables 5, 6 and 7, indicate that the welding process and plasma gas flow rate have significant effects on the geometric parameters of the bead.

In relation to the convexity index (CI = 100\*r/W), Silva et al. (2000) establishes that values close to 30% are desirable for the relation between the width (W) and reinforcement (r) of the weld bead. Figure 10 shows the convexity index of the weld bead for the PAW and PTA processes as a function of the plasma gas flow rate.

PTA Process

PAW Process

Figure 6 shows the external aspect of the beads where significant differences between them can be observed. The PTA process produced a better surface finish, better dilution, better

Figures 7 and 8 show cross-sections of the beads obtained using the two processes (PAW and PTA) where considerable differences in the penetration profile of the welds can be noted and Figure 9 shows the results for the geometric parameters of the beads, for the three levels of plasma gas flow rate tested in this study: 2.2; 2.4 and 3.0 l/min. On comparing the deposits obtained from the two processes it can be observed that the reinforcement and the penetration are always smaller in the PTA process (Figure 9). In the PTA process there was a significantly wider cord width, which is due to the use of a constrictor nozzle with a wider

The data shown in Figure 9 together with an analysis of the variance in Tables 5, 6 and 7, indicate that the welding process and plasma gas flow rate have significant effects on the

In relation to the convexity index (CI = 100\*r/W), Silva et al. (2000) establishes that values close to 30% are desirable for the relation between the width (W) and reinforcement (r) of the weld bead. Figure 10 shows the convexity index of the weld bead for the PAW and PTA

A cm/min l/min l/min l/min kg/h mm/º mm mm

mm m/min kg/h mm

A cm/min l/min l/min kg/h mm mm

160 20 2.2; 2.4; 3.0 10 2 1.4 4.8/30 10 2.4

> 1.2 3.0 1.4 3.2

160 20 2.2; 2.4; 3.0 10 1.4 10 2.4

Welding current Welding speed Plasma gas flow rate

Constrictor nozzle diameter/ convergence

Nozzle to workpiece distance

Wire diameter (tubular)

Constrictor nozzle diameter

Nozzle to workpiece distance

**3. Results and discussion 3.1 General characteristics**

geometric parameters of the bead.

processes as a function of the plasma gas flow rate.

wetting and wider width.

diameter.

Table 4. Welding variables and parameters.

Shield gas Carrier gas Feed rate

angle

Setback

Wire speed Deposition rate

Shield gas Feed rate

Setback

Welding current Welding speed Plasma gas flow rate Analysis of Figure 10 shows that for the three plasma gas flow rates tested the PTA process provided acceptable convexity of the weld beads (less than 30%), a highly desirable condition. In the case of the PAW process, the convexity index was acceptable only for low plasma gas flow rates.

The average values for the areas of the metal deposited varied for the two welding processes studied, as expected, due to the difference in the diameters of the constriction orifices used in each case and the material loss according to the efficiency of the deposition process.

Figure 11 shows that in the PTA process there was loss of material. Lin (1999) observed that losses occur mainly due to vaporization and also dispersion of the particles after making contact with the substrate.

Vergara (2005), reports that the carrier gas flow rate influences the dispersion of the particles. In many cases it is possible, at the end of the finishing operation, to observe unmolten powder particles adhered to the sides of the finish. On the other hand, when the deposition rate is very high (1.5 kg/h) in relation to the welding current (160 A) unmolten power can be seen spread over the substrate. Vergara [9] observed that the PTA process has a deposition efficiency of the order of 87% when a constrictor nozzle of 30º is used. Similar results have been reported by Davis (1993), who demonstrated a range of 85 to 95 % deposition yield for the PTA process.

The graph in Figure 12 shows the effect of the plasma gas flow rate on the degree of dilution using the wire Stellite 6, 1.2 mm tubular diameter. The results indicate that the dilution increases with the plasma gas flow rate possibly due to the greater pressure of the plasma jet. Similar results were found for the PTA process, with dilution values being lower than those achieved with the PAW process, as expected, due to the difference in the diameters of the constrictor orifice. Vergara (2005) reports that the diameter of the constrictor nozzle orifice has a considerable influence on the quality of the finish since it is directly related to the width and penetration of the weld bead produced. The data in Figure 12 together with the analysis of variance in Table 8 indicate that, in general, the welding process and the plasma gas flow rate significantly affect the dilution. Similar conclusions have been reported by Silvério (2003) for the alloy Stellite 1.

The good results obtained for the PTA process are associated with:

Fig. 6. Superficial aspect of Stellite 6 deposited by: a) PAW and b) PTA. Welding current = 160 A, Welding speed = 20 cm/min, Feed rate =1.4 kg/h, Plasma gas flow rate = 2.4 l/min.

Fig. 7. Cross-section of weld beads processed via PAW. Plasma gas flow rate: (a) 2.2 (l/min); (b) 2.4 (l/min); and (c) 3.0 (l/min)

(c)

Fig. 8. Cross-section of weld beads processed via PTA. Plasma gas flow rate: (a) 2.2 (l/min); (b) 2.4 (l/min); and (c) 3.0 (l/min).

(a) (b)

(c) Fig. 7. Cross-section of weld beads processed via PAW. Plasma gas flow rate: (a) 2.2 (l/min);

(a) (b)

Fig. 8. Cross-section of weld beads processed via PTA. Plasma gas flow rate: (a) 2.2 (l/min);

(b) 2.4 (l/min); and (c) 3.0 (l/min)

(b) 2.4 (l/min); and (c) 3.0 (l/min).

b) Reinforcement

Fig. 9. Effect of plasma gas flow rate on geometric parameters (Width, reinforcement, penetration).

Fig. 10. Effect of plasma gas flow rate on convexity index.

Obs.: Index of significance () = 5%

Table 5. Results of the analysis of variance for width.

Obs.: Index of significance () = 5%

Table 6. Results of analysis of variance for reinforcement.

#### Hardfacing by Plasma Transferred Arc Process 15

14 Arc Welding

17,7

2,2 2,4 3,0 2,2 2,4 3,0 **Plasma gas flow rate (l/min)**

Interaction 2.33 2 1.16 94.14 > 3.55

*Degrees of freedom*

*Degrees of freedom*

Interaction 0.098 2 0.049 15.45 > 3.55

21,4 19,4

*Average of squares*

*Average of squares*

PAW

PTA

*F observed F critical*

*F observed F critical*

40

28,6

Fig. 10. Effect of plasma gas flow rate on convexity index.

*squares*

*squares*

Residual 0.057 18 0.0032

Welding process 4.29 1 4.29 1353.78 Plasma gas flow rate 1.33 2 0.66 209.016

Residual 0.22 18 0.0124

Total 22.72 23

Table 5. Results of the analysis of variance for width.

Total 5.77 23

Table 6. Results of analysis of variance for reinforcement.

Welding process 17.85 1 17.85 1444.35 Plasma gas flow rate 2.316 2 1.16 93.67

*Source of variation Sum of*

Obs.: Index of significance () = 5%

Obs.: Index of significance () = 5%

*Source of variation Sum of*

**IC (%)**

36,6

Obs.: Index of significance () = 5%

Table 7. Results of analysis of variance for penetration.

Fig. 11. Area of material deposited in PAW and PTA processes.

Fig. 12. Effect of plasma gas flow rate on degree of dilution in PAW and PTA processes.

Obs.: Index of significance () = 5%

Table 8. Results of analysis of variance for dilution

#### **3.2 Microhardness and microstructure**

Figure 13 shows the typical microstructures of the solidification in the center of the weld bead. When a plasma gas flow rate of 2.2 l/min was used in the PAW and PTA processes the microstructure was more refined. For a plasma gas flow rate of 3.0 l/min for both welding processes the microstructure was less refined.

The microhardness profiles evaluated along the cross-section of the deposits are shown in Figures 14 and 15 for the PAW and PTA processes, respectively.

The data in Figure 14 together with the analysis of variance in Table 9, related to the PAW process, indicate that, in general, the plasma gas flow rate significantly affects the hardness. On the other hand, the data in Figure 15 together with the analysis of variance in Table 10, which relate to the PTA process, indicate that the plasma gas flow rate does not significantly affect the hardness. Deposits obtained with the PAW process have lower hardness values, which is to be expected given the less refined structures and higher degrees of dilution.

Obs.: Index of significance () = 5%

Table 9. Results of analysis of variance for average hardness of microstructure – PAW.

Interaction 25.93 2 12.96 297.4 > 3.55

Figure 13 shows the typical microstructures of the solidification in the center of the weld bead. When a plasma gas flow rate of 2.2 l/min was used in the PAW and PTA processes the microstructure was more refined. For a plasma gas flow rate of 3.0 l/min for both

The microhardness profiles evaluated along the cross-section of the deposits are shown in

The data in Figure 14 together with the analysis of variance in Table 9, related to the PAW process, indicate that, in general, the plasma gas flow rate significantly affects the hardness. On the other hand, the data in Figure 15 together with the analysis of variance in Table 10, which relate to the PTA process, indicate that the plasma gas flow rate does not significantly affect the hardness. Deposits obtained with the PAW process have lower hardness values, which is to be expected given the less refined structures and higher

> *Degrees of freedom*

Plasma gas flow rate 18214.93 2 9107.463 151.9637 > 3.2381

Table 9. Results of analysis of variance for average hardness of microstructure – PAW.

*Average of squares*

*Degrees of freedom*

*Average of squares*

*F observed F critical*

*F observed F critical*

*squares*

Residual 0.785 18 0.044

Total 1311.305 23

Table 8. Results of analysis of variance for dilution

welding processes the microstructure was less refined.

Figures 14 and 15 for the PAW and PTA processes, respectively.

*squares*

Residual 2337.341 39 59.93183

**3.2 Microhardness and microstructure**

*Source of variation Sum of*

Obs.: Index of significance () = 5%

Total 20552.27 41

Welding process 1102.43 1 1102.43 25289.88 Plasma gas flow rate 182.16 2 91.08 2089.39

*Source of variation Sum of*

Obs.: Index of significance () = 5%

degrees of dilution.

**c) Plasma gas flow rate = 2.2 (l/min)**

Fig. 13. Micrographs of the samples of Stellite 6 for the PAW and PTA processes. Centre of weld bead.

Fig. 14. Effect of plasma gas flow rate on hardness in PAW process.

Fig. 15. Effect of plasma gas flow rate on hardness in PTA process.

Obs.: Index of significance () = 5%

Table 10. Results of analysis of variance for average hardness of microstructure –– PTA.

It was verified that the PTA process generates a more refined microstructure and consequently greater hardness than the PAW process, as also observed by Silvério (2003).

### **4. Conclusions**

18 Arc Welding

PAW Process

VGP=3,0 (l/min) VGP=2,4 (l/min) VGP=2,2 (l/min)

PTA Process

VGP=3,0 (l/min) VGP=2,4 (l/min) VGP=2,2 (l/min)

0123456

0 0,5 1 1,5 2 2,5 3 3,5 4 4,5

Displacement from the surface (mm)

Fig. 15. Effect of plasma gas flow rate on hardness in PTA process.

Displacement from the surface (mm)

Fig. 14. Effect of plasma gas flow rate on hardness in PAW process.

0

0

100

200

300

Hardness (HV0,5)

400

500

600

100

200

300

Hardness (HV0,5)

400

500

600

Based on the experimental results obtained in this study the conclusions are as follows:

#### **5. References**

### **Fusion Welding with Indirect Electric Arc**

Rafael García1, Víctor-Hugo López1, Constantino Natividad2,

Ricardo-Rafael Ambriz3 and Melchor Salazar4

*1Instituto de Investigaciones Metalúrgicas-UMSNH, 2Facultad de Química-UNAM 3Instituto Politécnico Nacional CIITEC-IPN, 4Instituto Mexicano del Petróleo México*

#### **1. Introduction**

20 Arc Welding

Vergara, V. M. (2005). Inovação do equipamento e avaliação do processo plasma de arco

SIlva, C. R.; Ferraresi, V. A & Scotti, A. (2000). *A quality and cost approach for welding process*

LIN, J. A. (1999). *Simple model of powder catchment in coaxial laser cladding*. Optics & Laser

Davis, J. R. – Davis and Associates. (1993). *Hardfacing, Weld Cladding and Dissimilar Metal*

Silvério, R. B. & D'Oliveira , A.S. C. M. Revestimento de Liga a Base de Cobalto por PTA

<http://www.scielo.br/scielo.php?script=sci\_arttext&pid=S0100-

p. 537–9.

10.1590/S0100-73862000000300002.

ASM Metals Park. p. 699-828.

Fabricação, Uberlândia-MG, Maio. 2003.

Technology, 233-238.

transferido alimentado com pó (PTAP) para soldagem fora de posição: 2005. 174p. Doctoral Thesis, Mechanical Engineering Department - UFSC, Florianópolis. Hallen, H.; Lugscheider, E.; Ait-Mekideche, A. *Plasma transferred arc surfacing with high*

*deposition rates. In: Proceedings of conference on thermal spray coatings: properties, processes and applications*, Pittsburgh, USA, 4–10 May 1991. ASM International; 1992.

*selection.* J. Braz. Soc. Mech. Sci., Campinas, v. 22, n. 3. Available from

73862000000300002&lng=en&nrm=iso>. Accessed on 29 Nov. 2009. doi:

*Joining*. In: ASM Handbook – Welding, Brazing and Soldering, Vol. 6. 10th ed. OH:

com Alimentação de Pó e Arame. In: Congresso Brasileiro de Engenharía de

The indirect electric arc technique (IEA) is a welding process that was initially developed to weld aluminum metal matrix composites (MMCs) reinforced with high content of TiC particles. Later on, its use was extended to weld MMCs reinforced with low contents of SiC and Al2O3 particles and monolithic materials such as carbon steels, aluminum and aluminum alloys. This technique is based on using the gas metal arc welding process (GMAW). In this instance, however, fusion of the base metal is not realized by the direct contact between the electric arc and the work pieces. Instead, the application of the electric arc is on thin plates of feeding metal placed on top of the work pieces and aligned with the groove of the joint. The filler wire, fed in a spray transfer mode, forms a weld pool with the plates of feeding metal and the molten metal is instantaneously fed, at high temperature, into the groove of the joint. The heat input supplied with the molten metal melts the side walls of the work pieces enabling welding upon solidification. The IEA technique allows using feeding material with the same chemical composition of the base metal. It has been found that the microstructure obtained in the weld metal with this technique, in carbon steels, improves the resistance to stress corrosion in hydrogen sulfide. The IEA technique has proved to be effective in welding MMCs with low and high content of ceramic particles, aluminum and its alloys as well as carbon steels such as API X-65 employed for transport and storage of hydrocarbons.

The design of the IEA joint enables welding of plates, 12.5 mm thick, in a single welding pass with a reduced heat input and thereby a reduction in the thermal affection of the base metal. Trials to weld materials such as aluminum and MMCs with a thickness of 12.5 mm in one welding pass without joint preparation, i.e. square edges, resulted in deficient welds with partial penetration. Successful welding of these plates demands 3 or 4 welding passes using a single V joint design. Conversely, the use of the IEA technique with preheating of the joint led to welds with full penetration and without lack of fusion in the side walls in a sole welding pass. The multipass welding procedure required for the single V groove joint means a larger heat input which inevitably has an impact on the microstructure of the different regions of the welded joint and of course on its mechanical performance. A thermal balance of the IEA process revealed a larger thermal efficiency as compared to the traditional use of the GMAW welding process. This increase is ascribed to the fact that the electric arc is not openly exposed to the atmosphere. Instead, it is established in a hidden fashion within the groove formed by the feeding plates, reducing thus heat losses. This characteristic of the IEA technique has outstanding repercussions on the microstructure and mechanical properties of the welds. For example, degradation of the reinforcement during fusion welding of aluminum matrix composites reinforced with SiC particles is a common occurrence. Welding of this type of MMC with IEA did not show signs of reaction and a larger fraction of SiC particles were incorporated into the weld metal increasing the tensile properties of the welded joint. Also, in carbon steels, aluminum and its alloys the use of the IEA leads to a different solidification mode and grain refined microstructures. In particular, for an API X-65 this refinement has a profound positive effect in terms of sulfide stress cracking (SSC) behavior. Regarding heat treatable aluminum alloys, it is well known the overaging effect in the heat affected zone (HAZ) which weakens the strength of the alloy and predisposes failure in this region with a very low stress. The reduced heat input of the IEA technique also reduces the loss of strength in the HAZ of this type of alloys.

The main disadvantage of the IEA technique is that it leaves the residual feeding plates on top of the welded joint. This would be unacceptable in most of the applications. Thus, an additional step in the process demands removing these strips from the weld. To overcome this inconvenient, a modification of the design of the IEA joint was proposed and tested in heat treatable aluminum alloys. The use of the strips of feeding material on top of the work pieces was omitted and in the upper part of the work pieces a lash was machined, simulating the original feeding plates. The modified indirect electric arc (MIEA) technique drew similar results than the original IEA and when compared with the conventional single V groove joint, the behavior of the MIEA welds was better in both static and dynamic testing.

So far, the IEA technique and its evolution into the MIEA has emerged as an attractive alternative to weld a number of materials with peculiar microstructural characteristics that have a positive impact in the mechanical and stress corrosion cracking (SCC) behaviors. This chapter details a broad description of the process, emphasizing its advantages with respect to the conventional practice of fusion welding. An overview of the findings and benefits observed in different materials as well as the evolution of the original idea throughout ten years of research are provided.

#### **2. Overview of the IEA welding process**

The indirect electric arc (IEA) technique is a novel welding process that has been successfully used to join MMCs (Garcia et al, 2002a, 2002b, 2003). It is a variation of the metal inert gas (MIG) process in which fusion of the base metal is not caused by direct contact with the electric arc. Instead, the electric arc is established between the filler metal and feed metal, in plate form, placed over the base metal (Fig. 1), where the feed metal plates, base metal and filler metal all have similar chemical compositions. The plates are prepared with square edges and with a small single-V preparation with an angle of 45° in the upper part. The IEA technique allows using feeding material with the same chemical composition. The resultant droplets, in the form of a spray, produce a molten pool on the plates and the liquid is fed instantaneously at high temperature into the groove formed between the workpieces (Lu & Kou, 1989a, 1989b). The high temperature of the liquid metal melts the parent materials, producing the welded joint. Due to the increase of the thermal efficiency of the IEA process, complete penetration

traditional use of the GMAW welding process. This increase is ascribed to the fact that the electric arc is not openly exposed to the atmosphere. Instead, it is established in a hidden fashion within the groove formed by the feeding plates, reducing thus heat losses. This characteristic of the IEA technique has outstanding repercussions on the microstructure and mechanical properties of the welds. For example, degradation of the reinforcement during fusion welding of aluminum matrix composites reinforced with SiC particles is a common occurrence. Welding of this type of MMC with IEA did not show signs of reaction and a larger fraction of SiC particles were incorporated into the weld metal increasing the tensile properties of the welded joint. Also, in carbon steels, aluminum and its alloys the use of the IEA leads to a different solidification mode and grain refined microstructures. In particular, for an API X-65 this refinement has a profound positive effect in terms of sulfide stress cracking (SSC) behavior. Regarding heat treatable aluminum alloys, it is well known the overaging effect in the heat affected zone (HAZ) which weakens the strength of the alloy and predisposes failure in this region with a very low stress. The reduced heat input of the

IEA technique also reduces the loss of strength in the HAZ of this type of alloys.

testing.

years of research are provided.

**2. Overview of the IEA welding process**

The main disadvantage of the IEA technique is that it leaves the residual feeding plates on top of the welded joint. This would be unacceptable in most of the applications. Thus, an additional step in the process demands removing these strips from the weld. To overcome this inconvenient, a modification of the design of the IEA joint was proposed and tested in heat treatable aluminum alloys. The use of the strips of feeding material on top of the work pieces was omitted and in the upper part of the work pieces a lash was machined, simulating the original feeding plates. The modified indirect electric arc (MIEA) technique drew similar results than the original IEA and when compared with the conventional single V groove joint, the behavior of the MIEA welds was better in both static and dynamic

So far, the IEA technique and its evolution into the MIEA has emerged as an attractive alternative to weld a number of materials with peculiar microstructural characteristics that have a positive impact in the mechanical and stress corrosion cracking (SCC) behaviors. This chapter details a broad description of the process, emphasizing its advantages with respect to the conventional practice of fusion welding. An overview of the findings and benefits observed in different materials as well as the evolution of the original idea throughout ten

The indirect electric arc (IEA) technique is a novel welding process that has been successfully used to join MMCs (Garcia et al, 2002a, 2002b, 2003). It is a variation of the metal inert gas (MIG) process in which fusion of the base metal is not caused by direct contact with the electric arc. Instead, the electric arc is established between the filler metal and feed metal, in plate form, placed over the base metal (Fig. 1), where the feed metal plates, base metal and filler metal all have similar chemical compositions. The plates are prepared with square edges and with a small single-V preparation with an angle of 45° in the upper part. The IEA technique allows using feeding material with the same chemical composition. The resultant droplets, in the form of a spray, produce a molten pool on the plates and the liquid is fed instantaneously at high temperature into the groove formed between the workpieces (Lu & Kou, 1989a, 1989b). The high temperature of the liquid metal melts the parent materials, producing the welded joint. Due to the increase of the thermal efficiency of the IEA process, complete penetration and uniform weld beads, in a single pass, were obtained, as well as a reduction in the heat input and thereby a reduction in the thermal affection of the base metal compared with that provoked by the direct application of the electric arc (Garcia et al., 2002a, 2002b, 2003).

Fig. 1. IEA welding process: (a) General set-up (Garcia et al., 2002); (b) Set-up for application to 359 aluminum MMCs reinforced with 20%SiC (Garcia et al., 2007).

The fusion process of the base material is carried out by means of a liquid diffusion process, similar that the shown in Fig. 2a, where the liquid material diffuses through the grain boundaries, because these are areas of lower melting point than the matrix of the grains. This phenomenon is favored due to segregation of impurities and alloying elements in metals of high purity and alloying elements in alloys, respectively. Fig. 2b-c illustrates the difference between the indirect electric arc and the traditional electric arc welding processes. In the majority of the electric arc welding processes, the electric arc is established between the electrode and the base metal. The high energy developed by the electric arc is in direct contact with base metal and the forces generated in the weld pool affect the weld form and solidification mode. As illustrated in Fig. 2b, the temperature gradient within the weld pool induces a corresponding density gradient that enhances the flow and generates radial forces (RF) and circular forces (CF). Fig. 2b shows that when materials of a higher temperature and lower density are moving toward the bottom of the puddle, the buoyancy forces tend to rise up the liquid flow across the center of the pool. The flow moves radially outward, the molten metal is forced along the surface and then down the side of the weld pool toward the bottom (Domey et al., 1995). In the IEA technique, the electric arc is established within the feeding plates and as soon as the filler metal and feed metal are melted, liquid metal with low density is supplied into the joint geometry at high temperature. Thus, in the IEA process the buoyancy-driven flow is interrupted due to the presence of a deep groove and as a result, the radial forces instead of becoming circular forces, are transformed into "drag forces", a combined effect of the pressure exerted by the electric arc and the gravity action, drives the molten pool downwards into the groove formed by the workpieces, as can be seen in Fig. 2c. In addition, the hidden arc in the IEA method suggests that the loss of energy by radiation is suppressed and as a consequence, the efficiency of the electric arc in melting the feed metal is increased.

Along with the stir forces generated in the weld pool, the extent of supercooling also accounts for the mode of solidification. Although there are some exceptions, in electric arc welding epitaxial growth is a typical occurrence, wherein the first grains of the weld pool nucleate directly from randomly oriented grains in the HAZ and grow toward the greatest thermal gradient within the puddle (Domey et al., 1995). On the other hand, the solidification in the IEA welding method is different due to the distinct generation of forces and flow patterns in the weld pool. Fusion of the base metal is realized by the high energy of the melt supplied into the joint geometry. The improvement of the efficiency of the MIG welding process to about 95% using IEA may be ascribed to the better use of heat generated by the electric arc, which is established in a hidden form and the contact with the environment is minimum reducing thus heat losses.

Fig. 2. (a) Fusion process by liquid diffusion; Schematic of the MIG welding process using (b) direct electric arc and (c) indirect electric arc (Garcia et. al, 2002).

Fig. 3. Joint designs and typical geometry of the welds obtained; a) single V groove joint, b) IEA joint and c) MIEA joint.

The Fig. 3 shows different joint designs and dimensions as well as the typical geometries of the welds obtained by a) typical single V groove joint, b) IEA joint and c) MIEA joint. Whilst for the single V groove joint three or more welding passes are required to fill the groove, the IEA and MIEA joints only need one welding pass. The macrographs of the transverse weld profiles depict the geometries of the weld beads. Macroetching of the welds also revealed the HAZ. Roughly, it can be seen the larger thermal affection in the single V groove joint as compared to the IEA and MIEA welds.

#### **3. Applications of the IEA welding process**

In this section, an overview of the findings and benefits observed in different materials are provided. Between these materials are the MMCs, aluminium and its alloys and carbon steels.

#### **3.1 Composites**

24 Arc Welding

nucleate directly from randomly oriented grains in the HAZ and grow toward the greatest thermal gradient within the puddle (Domey et al., 1995). On the other hand, the solidification in the IEA welding method is different due to the distinct generation of forces and flow patterns in the weld pool. Fusion of the base metal is realized by the high energy of the melt supplied into the joint geometry. The improvement of the efficiency of the MIG welding process to about 95% using IEA may be ascribed to the better use of heat generated by the electric arc, which is established in a hidden form and the contact with the

(a) (b) (c)

(b) direct electric arc and (c) indirect electric arc (Garcia et. al, 2002).

Fig. 2. (a) Fusion process by liquid diffusion; Schematic of the MIG welding process using

Fig. 3. Joint designs and typical geometry of the welds obtained; a) single V groove joint,

environment is minimum reducing thus heat losses.

Base Metal Melted

a)

b)

c)

Filler Metal Melted Base Metal Partially Melted

b) IEA joint and c) MIEA joint.

The development of MMCs was a breakthrough in materials technology in the 80´s. Acceptation and use of new materials rely on their readiness to be joined. Sorting out the challenges of incorporating ceramic reinforcements into molten metals and alloys was not enough for spreading the use of MMCs. A major problem was also encountered when trying to join this type of materials with conventional fusion welding processes. Exposure of the Al/SiC-type composite to temperature above the liquidus of the aluminium alloys, as typically experienced in welding, results in a severe lose of mechanical properties due to the formation of brittle and hygroscopic aluminium-carbon compounds, mainly aluminium carbide. In addition, fusion welding processes produce a weld pool that has poor fluidity and solidifies with large volumes of porosity in the weld, because of the realization of hydrogen from the melted aluminium powder, which is used to make many MMCs (Ahearn et al. 1982). A reduction in the porosity in welds deposited by gas tungsten arc welding (GTAW) was achieved by previous vacuum degassing for long periods of time before welding. Nevertheless, both Al4C3 and Al-Si eutectic were detected in degassed composites. A number of authors (Cola & Lundin, 1989; Devletian 1987; Fukumoto & Linert, 1993; J Ahearn, et al, 1982, Ellis, et al, 1995; Urena et al, 2000, Lundin et al, 1989, Lienert, et al, 1993) reported that Al4C3 is always formed in the weld metal in MMCs reinforced with SiC no matter which of the fusion welding processes is employed to weld the MMCs (laser, electron beam, TIG, MIG and so on), and the formation of this compound occurs according to Eq. (1).

$$\text{GSiC}\_{\text{(s)}} + \text{4Al}\_{\text{(l)}} \rightarrow \text{Al}\_{\text{4}}\text{C}\_{\text{3}} + \text{3Si}\_{\text{(s)}} \tag{1}$$

This reaction is not reversible and the Al4C3 is formed as plates in the microstructure. The presence of the plates has two deleterious effects. First, the material becomes extremely brittle, and second, it becomes very prone to corrosion in presence of water, leading to the release of acetylene gas. In an extreme case, this has led to total corrosion of the weld within a few days. In response to the problematic issue of welding MMCs, the idea of the indirect electric arc was conceived (Garcia et al 2002) with the metal inert gas (MIG) welding process in order to overcome the difficulties of welding MMCs. The concept is based on the fact that experimental measurements indicate that the temperature of the droplets in the MIG welding process with spray transfer is between 2000 to 2327 °C for aluminium and its alloys according to (Lu & Kou, 1989, Kim et al, 1991). If a molten metal with a large overheating is casted into a "mould" shaped by the parent materials to be welded, the sensible heat of the weld pool formed is sufficient to melt the side walls (the matrix in MMCs) whilst instantly filling the groove, yielding the welded joint upon freezing. The indirect application of the electric arc reduces the degradation of the ceramic but still the temperature of the molten metal is large to induce spontaneous and instantaneous wetting so that continuity is seen between weld metal and the matrix of the composites when the content of reinforcement is large (Garcia et al, 2002, 2003) and significant incorporation of particles into the weld metal occurs for composites with low fraction of reinforcement (Garcia et al, 2002, 2007).

The MIG welding process with IEA is a novel fusion welding method, which was developed to join MMCs with a reduced HAZ in the base metal. In the IEA welding method, the fusion of the base metal is not caused by direct contact with the electric arc, instead, the electric arc is established between the solid electrode and a plate of the same base metal, placed over the parent material. The resultant droplets, in the form of a spray, produce a molten pool on the plates and the liquid is fed instantaneously at high temperature into the groove formed by the base metal; Fig. 4 depicts the experimental setup in a dissimilar join. Experimental measurement indicate that the droplets temperature in the MIG welding process with spray transfer is between 2000 to 2327 °C for aluminium and its alloys according to (Lu & Kou, 1989, Kim et al, 1991). As a result of its elevated temperature, the liquid metal melts the parent materials (the matrix in MMCs), yielding the welded joint upon freezing.

Profiles of MMCs welds using IEA are shown in Fig.4. Irrespective of the reinforcement content (high, medium or null), full penetration was attained in one welding pass and uniform weld profiles are obtained with little fusion of the base materials and a minimized heat input. The contour of the weld at the top depicts the configuration of the electric arc, which does not impinge on the surface of the parent plates; rather, it strikes inside the channel formed by them, as illustrated previously in Fig.1. Thus, during welding the electric arc is hidden and the typical flashing and sputtering of the normal MIG welding process is no longer observed when welding is carried out in any kind of material. It is well known that in order to weld 9, 10 mm thick MMCs and 12.5 mm thick aluminium plates, conventional MIG welding practice demands more than one welding pass with low travel speed, which leads to a large HAZ. It is worthy bearing in mind that most of the attempts to weld MMCs have been performed rather in thin sections or depositing bead on plate welds.

It has been reported during MMCs welding by different welding processes, that the high energy developed by the electric arc produces a wide HAZ accompanied by dissociation of the ceramic particles and the formation of hygroscopic compounds (Al4C3). This was confirmed by (Garcia et al, 2007) when welding an A359/SiC/20p commercial composite. Fig. 5a shows jagged SiC particles within the weld metal, this feature was not seen when the composite was welded with IEA. Fig 5b shows also the particles incorporated into the weld metal but they retain their initial angularity meaning that significant degradation (according to the resolution of the optical microscope) did not occur during welding with IEA. Tensile testing of the welded joints drew a tensile strength of 234 MPa for the IEA weld (one welding pass) as compared to 209 MPa for the weld with direct application of the electric arc (three welding passes). This behaviour is related to a larger incorporation of SiC particles into the weld metal and the reduced porosity for the IEA weld. The authors stated that the degradation of the SiC particles observed in the plain weld played a minor role during mechanical testing.

casted into a "mould" shaped by the parent materials to be welded, the sensible heat of the weld pool formed is sufficient to melt the side walls (the matrix in MMCs) whilst instantly filling the groove, yielding the welded joint upon freezing. The indirect application of the electric arc reduces the degradation of the ceramic but still the temperature of the molten metal is large to induce spontaneous and instantaneous wetting so that continuity is seen between weld metal and the matrix of the composites when the content of reinforcement is large (Garcia et al, 2002, 2003) and significant incorporation of particles into the weld metal

The MIG welding process with IEA is a novel fusion welding method, which was developed to join MMCs with a reduced HAZ in the base metal. In the IEA welding method, the fusion of the base metal is not caused by direct contact with the electric arc, instead, the electric arc is established between the solid electrode and a plate of the same base metal, placed over the parent material. The resultant droplets, in the form of a spray, produce a molten pool on the plates and the liquid is fed instantaneously at high temperature into the groove formed by the base metal; Fig. 4 depicts the experimental setup in a dissimilar join. Experimental measurement indicate that the droplets temperature in the MIG welding process with spray transfer is between 2000 to 2327 °C for aluminium and its alloys according to (Lu & Kou, 1989, Kim et al, 1991). As a result of its elevated temperature, the liquid metal melts the parent materials (the matrix in MMCs), yielding

Profiles of MMCs welds using IEA are shown in Fig.4. Irrespective of the reinforcement content (high, medium or null), full penetration was attained in one welding pass and uniform weld profiles are obtained with little fusion of the base materials and a minimized heat input. The contour of the weld at the top depicts the configuration of the electric arc, which does not impinge on the surface of the parent plates; rather, it strikes inside the channel formed by them, as illustrated previously in Fig.1. Thus, during welding the electric arc is hidden and the typical flashing and sputtering of the normal MIG welding process is no longer observed when welding is carried out in any kind of material. It is well known that in order to weld 9, 10 mm thick MMCs and 12.5 mm thick aluminium plates, conventional MIG welding practice demands more than one welding pass with low travel speed, which leads to a large HAZ. It is worthy bearing in mind that most of the attempts to weld MMCs have been performed rather in thin sections or

It has been reported during MMCs welding by different welding processes, that the high energy developed by the electric arc produces a wide HAZ accompanied by dissociation of the ceramic particles and the formation of hygroscopic compounds (Al4C3). This was confirmed by (Garcia et al, 2007) when welding an A359/SiC/20p commercial composite. Fig. 5a shows jagged SiC particles within the weld metal, this feature was not seen when the composite was welded with IEA. Fig 5b shows also the particles incorporated into the weld metal but they retain their initial angularity meaning that significant degradation (according to the resolution of the optical microscope) did not occur during welding with IEA. Tensile testing of the welded joints drew a tensile strength of 234 MPa for the IEA weld (one welding pass) as compared to 209 MPa for the weld with direct application of the electric arc (three welding passes). This behaviour is related to a larger incorporation of SiC particles into the weld metal and the reduced porosity for the IEA weld. The authors stated that the degradation of the SiC particles observed in the plain weld played a minor role during

occurs for composites with low fraction of reinforcement (Garcia et al, 2002, 2007).

the welded joint upon freezing.

depositing bead on plate welds.

mechanical testing.

Fig. 4. Weld profiles obtained with the MIG-IEA technique in different MMCs. a) Al-1010/TiC/50p (Garcia et al, 2003), b) Al-6061/Al2O3/20p (Garcia et al, 2002), c) A359/SiC/20p and d) dissimilar joint.

Fig. 5. Effect of the a) direct and b) indirect application of the electric arc on the SiC particles during welding an A359/SiC/20p commercial composite (Garcia et al, 2007).

#### **3.2 Carbon steels**

Pipelines of low carbon steel welded by electric arc have been used for many years and are widely used in the petroleum industry. However, frequent failures during operation over the years have prompted several studies of the design, construction, operation and maintenance of equipment and metallic structures used in this industry (Craig, 1997). Oil and gas from Mexico contain high concentrations of H2S and CO2, these constituents induce failures. In the oil industry, the welding processes commonly used for joining pipeline are electrical resistance (ERW) and submerged arc welding (SAW). With these processes, wide HAZs are created and as a result, a high probability of cracking due to the heterogeneous weldment exists. Studies of weld bead failures have demonstrated that these occur mainly in the HAZ because of the variation in microstructure (grain growth), residual stresses and a higher susceptibility to embrittlement by hydrogen. These microstructures are produced by the thermal gradients experienced in the joint during welding. Therefore, it is very important to develop welding processes with a narrow HAZs, this is possible with the IEA process. The welding process has an important influence in the SSC susceptibility of materials. The different welding processes promote different changes in the microstructure of the welded zone; these changes affect the SCC resistance and the yield strength of the pipeline steel.

A comparative study of SSC resistance between IEA, SAW and MIG was carried out (Natividad et. al., 2007) through slow strain rate tests (SSRT), electrochemical tests and hydrogen permeation measurements. The base metal used was API grade X-65 pipeline steel. Cylindrical tensile specimens with a gauge length of 25 mm and gauge diameter of 2.50 mm were machined from the pipeline perpendicular to the rolling direction. The specimens were subjected to conventional slow strain rate tests in air (as an inert environment) and in the standard NACE solution (5% NaCl, 0.5% acetic acid, saturated with hydrogen sulphide (H2S)) at a strain rate of 1x10-6 s-1 at room temperature (25°C) and at 37°C and 50°C. All of the tests were performed at the open circuit potential (OCP). The loss in ductility was assessed in terms of the percentage reduction in cross-sectional area (% RA), as follows

$$\%RA = \frac{A\_1 - A\_f}{A\_f} \ge 100\tag{2}$$

where Ai and Af are the initial and final cross-sectional areas, respectively. The index of susceptibility to SSC (*ISSC*) was calculated as follows

$$I\_{SSC} = \frac{\%RA\_{AIR} - \%RA\_{NACE}}{\%RA\_{AIR}} \tag{3}$$

where % RAAIR and % RANACE are the percentage reduction in area values in air and in the H2S-saturated NACE solution, respectively. An *ISSC* value close to unity indicates high susceptibility towards SSC whereas a value close to zero indicates immunity. The fracture surfaces were then examined using scanning electron microscopy (SEM).

Fig. 6 shows the macro and microstructures of the weldments obtained by the three welding processes. This figure shows clearly the different zones: base metal (BM), weld bead (WB), HAZ and fusion zone (FZ). Full penetration and a narrow HAZ are observed in microstructure obtained by IEA process. The weld bead and HAZ is very different from that obtained by both SAW and MIG processes. In general, the microstructure obtained with the IEA welding process is more homogeneous than the obtained by SAW and MIG processes. The corrosion and the SSC susceptibility of the welds are affected by the differences in composition, microstructure, and electrochemical potentials among the different zones. A lower electrochemical potential of the weld bead is commonly related to composition, microstructure and distribution of inclusions (Dawson et. al., 1997). Similarly, in a study performed by Turnbull and Nimmo, about SCC susceptibility (Turnbull & Nimmo, 2005), a direct relation to OCPs with mechanical properties like microstructure or hardness of phases was reported.

HAZs are created and as a result, a high probability of cracking due to the heterogeneous weldment exists. Studies of weld bead failures have demonstrated that these occur mainly in the HAZ because of the variation in microstructure (grain growth), residual stresses and a higher susceptibility to embrittlement by hydrogen. These microstructures are produced by the thermal gradients experienced in the joint during welding. Therefore, it is very important to develop welding processes with a narrow HAZs, this is possible with the IEA process. The welding process has an important influence in the SSC susceptibility of materials. The different welding processes promote different changes in the microstructure of the welded zone; these changes affect the SCC resistance and the yield strength of the

A comparative study of SSC resistance between IEA, SAW and MIG was carried out (Natividad et. al., 2007) through slow strain rate tests (SSRT), electrochemical tests and hydrogen permeation measurements. The base metal used was API grade X-65 pipeline steel. Cylindrical tensile specimens with a gauge length of 25 mm and gauge diameter of 2.50 mm were machined from the pipeline perpendicular to the rolling direction. The specimens were subjected to conventional slow strain rate tests in air (as an inert environment) and in the standard NACE solution (5% NaCl, 0.5% acetic acid, saturated with hydrogen sulphide (H2S)) at a strain rate of 1x10-6 s-1 at room temperature (25°C) and at 37°C and 50°C. All of the tests were performed at the open circuit potential (OCP). The loss in ductility was assessed in terms

> <sup>1</sup> % 100 *<sup>f</sup> f A A RA x A*

where Ai and Af are the initial and final cross-sectional areas, respectively. The index of

% % %

where % RAAIR and % RANACE are the percentage reduction in area values in air and in the H2S-saturated NACE solution, respectively. An *ISSC* value close to unity indicates high susceptibility towards SSC whereas a value close to zero indicates immunity. The fracture

Fig. 6 shows the macro and microstructures of the weldments obtained by the three welding processes. This figure shows clearly the different zones: base metal (BM), weld bead (WB), HAZ and fusion zone (FZ). Full penetration and a narrow HAZ are observed in microstructure obtained by IEA process. The weld bead and HAZ is very different from that obtained by both SAW and MIG processes. In general, the microstructure obtained with the IEA welding process is more homogeneous than the obtained by SAW and MIG processes. The corrosion and the SSC susceptibility of the welds are affected by the differences in composition, microstructure, and electrochemical potentials among the different zones. A lower electrochemical potential of the weld bead is commonly related to composition, microstructure and distribution of inclusions (Dawson et. al., 1997). Similarly, in a study performed by Turnbull and Nimmo, about SCC susceptibility (Turnbull & Nimmo, 2005), a direct relation to OCPs with mechanical properties like microstructure or hardness of phases

*RA RA <sup>I</sup>*

*AIR NACE*

*AIR*

*RA*

(2)

(3)

of the percentage reduction in cross-sectional area (% RA), as follows

*SSC*

surfaces were then examined using scanning electron microscopy (SEM).

susceptibility to SSC (*ISSC*) was calculated as follows

pipeline steel.

was reported.

The limit of hardness recommended for avoiding cracking in a sour environment is 22 Rc (248 Hv) (NACE/ISO, 2009). Although susceptibility to SSC generally increases with increasing hardness, some microstructures are more susceptible to cracking than others at the same hardness. Fig. 7 shows hardness measurements obtained from a transverse section of the weld for each welding process and it is observed that the values of hardness in the weld bead made by the SAW process are the lowest. For the MIG process, there is no significant difference between the HAZ and the WB hardness values. On the other hand, for the IEA process, the hardness value of the HAZ decreases by nearly 35 HV with respect to weld bead values. These values are within the recommended limits to avoid the fracture and cracking of the weld bead.

Fig. 6. Macro and microstructures of the weldments obtained by (a–b) SAW, (c–d) MIG and (e–f) IEA, (Natividad et al., 2007).

Fig. 7. Hardness values in weldments obtained by the different welding processes. (Natividad et al., 2007).

Fig. 8 shows the *ISCC* measurements on the three weldments at the three bath temperatures. The welded joint obtained by SAW shows the highest *ISCC* values, which was correlated with the hydrogen permeation results. Similarly, the *ISCC* measurements in the IEA case indicate a lower concentration of both the trap sites in the metal and the susceptibility to hydrogen embrittlement, because here the most hydrogen flux passed to the anodic cell side. Although the high electrochemical activity generates a higher atomic hydrogen concentration, a smaller concentration of this atomic hydrogen was trapped in the bulk. In addition, the *ISCC* results were affected by the change of welded microstructure, where the IEA presented a refined higher concentration of bainite compared with the grain coarse ferrite+pearlite microstructure obtained by the SAW process as reported before (Turnbull & Nimmo, 2005), and the consequent change in hardness due to the modified grain size during welding process (Omweg et al, 2003). Some evidence of the hydrogen diffusion effect into the welded joints is presented in the SCC fractographs illustrated in Fig. 9. The SAW weldments (Fig. 9b) show a more brittle fracture than the MIG and IEA weldments. The IEA weldment presented a less brittle fracture (Fig. 9a) and the *ISCC* values show this behaviour. However, the three welded joints do not show a completely brittle behaviour, but in general, the fracture behaviour was closer to a quasi-cleavage fracture. The hydrogen diffusion was low in quantity and low in permanence time into the electrolyte generating hydrogen embrittlement, but there is evidence of the hydrogen damage to the weldments. Additionally in this work, the IEA material was the least susceptible to hydrogen embrittlement damage, of course, the SCC resistance was higher, and was related to the lowest OCP activity promoted by the change in microstructure of the weldment.

**MIG**

**SAW**

**IEA**

**0 1 2 3 4 5 6 7 8 9 10 11 12 13 14**

**Distance (mm)**

Fig. 8 shows the *ISCC* measurements on the three weldments at the three bath temperatures. The welded joint obtained by SAW shows the highest *ISCC* values, which was correlated with the hydrogen permeation results. Similarly, the *ISCC* measurements in the IEA case indicate a lower concentration of both the trap sites in the metal and the susceptibility to hydrogen embrittlement, because here the most hydrogen flux passed to the anodic cell side. Although the high electrochemical activity generates a higher atomic hydrogen concentration, a smaller concentration of this atomic hydrogen was trapped in the bulk. In addition, the *ISCC* results were affected by the change of welded microstructure, where the IEA presented a refined higher concentration of bainite compared with the grain coarse ferrite+pearlite microstructure obtained by the SAW process as reported before (Turnbull & Nimmo, 2005), and the consequent change in hardness due to the modified grain size during welding process (Omweg et al, 2003). Some evidence of the hydrogen diffusion effect into the welded joints is presented in the SCC fractographs illustrated in Fig. 9. The SAW weldments (Fig. 9b) show a more brittle fracture than the MIG and IEA weldments. The IEA weldment presented a less brittle fracture (Fig. 9a) and the *ISCC* values show this behaviour. However, the three welded joints do not show a completely brittle behaviour, but in general, the fracture behaviour was closer to a quasi-cleavage fracture. The hydrogen diffusion was low in quantity and low in permanence time into the electrolyte generating hydrogen embrittlement, but there is evidence of the hydrogen damage to the weldments. Additionally in this work, the IEA material was the least susceptible to hydrogen embrittlement damage, of course, the SCC resistance was higher, and was related to the

BM HAZ HAZ BM

Weld Metal

Fig. 7. Hardness values in weldments obtained by the different welding processes.

lowest OCP activity promoted by the change in microstructure of the weldment.

**Hardness**

(Natividad et al., 2007).

**70**

**105**

**140**

**175**

**210**

**245**

 **(Hv10)**

Fig. 8. Variation of *Iscc* as function of temperature on the weldments (Natividad et al., 2007).

With the features of the IEA process, a better SSC resistance at 25 °C was obtained in comparison to the SAW and MIG processes. For specimens obtained by the IEA process subjected to SSR tests, the failure occurred on the base metal and for specimens produced by SAW and MIG processes it occurred in the weld bead and HAZ respectively. At 37 °C and 50 °C, the SCC resistance of all processes show a similar behaviour (failed in the base metal). The higher atomic hydrogen permeation flux presented by the IEA process was promoted by the ferrite phase from the base metal, which is comparable to the fracture which occurred from SSRT.

A significant increase in the corrosion current (more positive OCP) was presented by the IEA material, and was attributed to the greatest galvanic cell formation between the welded and base metal, but this material shows superior resistance to the SCC susceptibility than that for the MIG and SAW processes. The hydrogen permeation results and the difference in microstructure presented by the IEA process corroborate this behaviour.

Fig. 9. SEM images of the SCC fractographies of (a) IEA weldment material and (b) SAW weldment material tested at 25 °C into the NACE solution saturated with H2S, (Natividad et al., 2007).

#### **3.3 Aluminium and Al-alloys**

Aluminium and its alloys are important engineering materials with a vast number of applications as structural and functional components. The shiny and attractive appearance of these materials is due to the native oxide layer that always envelops the bulk of the material. The aluminium oxide layer is highly compact and this characteristic prevents further thickening of the oxide and permeation of aggressive media, making of Al and its alloys corrosion resistant materials in a number of environments. Besides, their lightness means a large strength/weight ratio for heat treatable Al-alloys (Davies, 1993 & Heinz et al, 1990). The high technological relevance of Al and its alloys demands quite often joining operations. The first barrier to overcome for successful joining is the native oxide layer which avoids coalescence between faying surfaces. In order to succeed fusion welding with the electric arc, direct current and reverse polarity are used to generate an ionic striking effect that dissolves the oxide and enables mixing between molten metals. Another aspect to take into consideration is the high solubility of hydrogen in aluminium melts. If care is not taken in preventing sources of hydrogen during welding, this effect may lead to welds with a large level of porosity in the weld metal. Besides, Al-alloys are prone to solidification and liquation cracking owing to their relatively high thermal expansion, large change in volume upon freezing as well as wide solidification temperature range (Gittos et al, 1981; Enjo & Kuroda, 1982; Kerr & Katoh, 1987; Miyazaki et al, 1990; Malin, 1995; Huang & Kou, 2004). All the above mentioned defects may have a profound impact on the mechanical performance of the welded joint and therefore they must be prevented. Furthermore, exposure of heat treatable Al-alloys to the welding thermal cycles gives rise to a soft zone in the HAZ caused by overaging where the mechanical strength may decrease up to 50% (Malin, 1995). Partial or almost full recovery of the mechanical properties can be achieved by post weld heat treating (solutioning, quenching and aging) but this procedure might be restricted in many instances and the welded joint must be on service in the as-welded condition. On this context, contributions aimed to improve the mechanical performance of aluminium alloys are desirable and valuable.

Fig. 10. Comparison of the weld profiles obtained with the direct (a-b) and indirect (c-d) application of the electric arc with preheating of the joint (Garcia et al, 2007).

Aluminium and its alloys are important engineering materials with a vast number of applications as structural and functional components. The shiny and attractive appearance of these materials is due to the native oxide layer that always envelops the bulk of the material. The aluminium oxide layer is highly compact and this characteristic prevents further thickening of the oxide and permeation of aggressive media, making of Al and its alloys corrosion resistant materials in a number of environments. Besides, their lightness means a large strength/weight ratio for heat treatable Al-alloys (Davies, 1993 & Heinz et al, 1990). The high technological relevance of Al and its alloys demands quite often joining operations. The first barrier to overcome for successful joining is the native oxide layer which avoids coalescence between faying surfaces. In order to succeed fusion welding with the electric arc, direct current and reverse polarity are used to generate an ionic striking effect that dissolves the oxide and enables mixing between molten metals. Another aspect to take into consideration is the high solubility of hydrogen in aluminium melts. If care is not taken in preventing sources of hydrogen during welding, this effect may lead to welds with a large level of porosity in the weld metal. Besides, Al-alloys are prone to solidification and liquation cracking owing to their relatively high thermal expansion, large change in volume upon freezing as well as wide solidification temperature range (Gittos et al, 1981; Enjo & Kuroda, 1982; Kerr & Katoh, 1987; Miyazaki et al, 1990; Malin, 1995; Huang & Kou, 2004). All the above mentioned defects may have a profound impact on the mechanical performance of the welded joint and therefore they must be prevented. Furthermore, exposure of heat treatable Al-alloys to the welding thermal cycles gives rise to a soft zone in the HAZ caused by overaging where the mechanical strength may decrease up to 50% (Malin, 1995). Partial or almost full recovery of the mechanical properties can be achieved by post weld heat treating (solutioning, quenching and aging) but this procedure might be restricted in many instances and the welded joint must be on service in the as-welded condition. On this context, contributions aimed to improve the mechanical performance of aluminium alloys are desirable and valuable.

Fig. 10. Comparison of the weld profiles obtained with the direct (a-b) and indirect (c-d)

**Al**

**d) 100°C**

**b) 100°C**

**10 mm**

**2024**

**10 mm**

application of the electric arc with preheating of the joint (Garcia et al, 2007).

**10 mm**

 **10 mm**

**3.3 Aluminium and Al-alloys**

**c) 50°C**

**a) 50°C**

The IEA technique developed by (Garcia et al, 2002) was applied to study the feasibility of welding Al and its alloys by using an ER4043 filler wire. Preliminary welding trials (Garcia et al, 2007), in plates of aluminium of commercial purity, revealed that the IEA technique has the capability of joining plates of 12.5 mm in thickness in a single welding pass with full penetration by preheating the plates to 50 or 100°C as shown in Fig. 10. Conversely, with the same preparation of the parent plates but without the feeding plates, welds with partial penetration are obtained as seen in Fig. 10. Needless to say, this problem shows why traditionally welding of thick sections is performed by a multi-pass procedure either with a single or double V preparation of the plates.

Microstructural characterization of the welds disclosed for the welds with direct application of the electric arc the typical epitaxial and columnar grain growth from the partially melted grains of the base metal whereas for the IEA welds, partially melted grains of the base metal were found trapped within the weld metal as shown in Fig. 11. The mechanism of this phenomenon was previously mentioned in section 2. Dragging and survival of partially melted grains modified the mode of solidification in the weld metal by blocking columnar growth. Thermal analysis of the IEA technique showed that heat losses by radiation and convection are reduced since the electric arc is established in a hidden fashion within the channel formed by the feeding plates. The authors stated that the low ionization potential of the aluminium vapour, 5.986 eV, plays also an important role in the IEA technique.

Fig. 11. Optical micrographs of the base metal/weld interface: (a) in the transversal and (b) longitudinal axes of a MIG-IEA weld (Garcia et al, 2007).

The preliminary findings suggested that if the number of welding passes are reduced, so it is the heat input. This advantage means also a reduction in the thermal affection of the base metal and therefore a minimization in the loss of strength in heat treatable aluminium alloys. To prove this, plates of an Al-6061-T6 were welded with the IEA technique and the results were compared with a single V joint (Ambriz et al, 2009). Table 1 shows the results of the mechanical properties under tension of the welded joints in the as-welded condition. At first sight, it is clear that the mechanical properties are lower than the base metal independently of the joint design employed. The worst mechanical properties were, however, exhibited by the single V-groove joint, which has approximately 37.4 and 61% mechanical efficiency with respect to base material and filler wire respectively. In comparison to the single V groove joint, the IEA joints exhibited a notable increase in the mechanical efficiency with values ranging between 48 to 55%. Nevertheless, those efficiencies are still below the permissible limit of 57% for the 6061-T6 alloy according to the ASME Pressure Vessel Code, Section VIII. Irrespective of the joint design and preheating condition, the mechanical failure under tensile testing occurred in the base metal, at different distances from the fusion line, depending on joint design and preheating condition. The nearest failure from the fusion line occurred for the single V joint.

Table 1. Mechanical properties of the weldments in the as-welded condition (Ambriz et al, 2009).

Fig. 12. Microhardness profiles for; a) the single V joint and b) the IEA joint preheated to 100°C (Ambriz et al, 2009).

Fig. 12 shows microhardness profiles along the cross sections of the welds, they corroborate the failure zone with the variation of mechanical properties, as a result of the thermal cycle of welding. The plots of these figures reveal, in each weld, a soft zone which matches well with the failure zone of the tensile specimens. This behavior, which nevertheless was expected, agrees with others (Gutierrez et al, 1996; Liu et al, 1991; Myhr et al, 2004; Shelwatkar, 2002; Kostrivas & Lippold, 2000). Worth to point out is the fact that the depth, in terms of microhardness, and width of the soft zone is related to the joint design and preheating temperature. These features are obviously reflected on the tensile strength of the welds. Note that although the IEA and single V groove are similarly "soft", the later fails

mechanical efficiency with values ranging between 48 to 55%. Nevertheless, those efficiencies are still below the permissible limit of 57% for the 6061-T6 alloy according to the ASME Pressure Vessel Code, Section VIII. Irrespective of the joint design and preheating condition, the mechanical failure under tensile testing occurred in the base metal, at different distances from the fusion line, depending on joint design and preheating condition.

> Elongation (%)

Failure from fusion line (mm)

Efficiency from base metal (%)

Efficiency from filler metal (%)

The nearest failure from the fusion line occurred for the single V joint.

Yield strength (MPa)

ER-4043 -- 190 164 -- -- -- -- 6061-T6 -- 328 300 14 -- -- -- DEA 25 116 57 17.6 3 35.36 61.05 IEA 50 148.9 71.6 14 11-13 45.39 78.36 IEA 100 172.1 76.8 14.8 13-15 52.46 90.57 IEA 150 168 86.7 15.2 13-16 51.21 88.42 Table 1. Mechanical properties of the weldments in the as-welded condition (Ambriz et al,

Fig. 12. Microhardness profiles for; a) the single V joint and b) the IEA joint preheated to

Fig. 12 shows microhardness profiles along the cross sections of the welds, they corroborate the failure zone with the variation of mechanical properties, as a result of the thermal cycle of welding. The plots of these figures reveal, in each weld, a soft zone which matches well with the failure zone of the tensile specimens. This behavior, which nevertheless was expected, agrees with others (Gutierrez et al, 1996; Liu et al, 1991; Myhr et al, 2004; Shelwatkar, 2002; Kostrivas & Lippold, 2000). Worth to point out is the fact that the depth, in terms of microhardness, and width of the soft zone is related to the joint design and preheating temperature. These features are obviously reflected on the tensile strength of the welds. Note that although the IEA and single V groove are similarly "soft", the later fails

a) b)

Ultimate strength (MPa)

Joint type

2009).

100°C (Ambriz et al, 2009).

Preheating temperature (°C)

with a lower stress due to the relaxation of the dislocations as a result of repeated exposure to heat. Another important aspect revealed by the microhardness profiles is the increase in hardness within the weld metal for the IEA joint as compared to the single V groove joint. This trend is related to the dilution rates for each joint, the larger the dilution the largest incorporation of alloying elements into the weld metal, and to the solidification mode inherent to every joint design. Whilst the single V groove joints exhibits the typical epitaxial and competitive-columnar grain growth, the IEA joints solidified with a grain refined microstructure.

#### **4. Modified indirect electric arc (MIEA)**

The main disadvantage of the IEA technique is that it leaves the residual feeding plates on top of the welded joint. This would be unacceptable in most of the applications. Thus, an additional step in the process demands removing these strips from the weld. To overcome this inconvenient, a modification of the design of the IEA joint was proposed and tested in heat treatable aluminum alloys. The use of the strips of feeding material on top of the work pieces was omitted and in the upper part of the work pieces a lash was machined, simulating the original feeding plates. This evolution was named MIEA (Ambriz at al., 2006, Ambriz et al. 2008). The MIEA provided several advantages with respect to the traditional arc fusion welding process, for instance:

MIEA welding technique employs the same equipment that is required to weld by GMAW. General dimensions and the schematic representation of the MIEA joint along with the weld bead geometry obtained were shown in Fig. 3c. The dissipation of heat in this case is quite similar to the IEA joint so that large thermal efficiency is also obtained.

#### **4.1 Microstructure in aluminum alloys welds**

In a fusion welding process, the heat input produces a fusion-solidification phenomenon, which is different to that obtained in the solidification of an ingot. (i) In an ingot, solidification begins with heterogeneous nucleation at the chill zone meanwhile in a weld pool the liquid metal partially wets the grains of the parent metal and epitaxial growth takes place from the partially melted grains of the parent metal (Davies et al., 1975). (ii) The rate of solidification in a weld pool, which depends on the traveling speed as well as the welding process, is by far faster than in an ingot. (iii) The macroscopic profile of the solid/liquid interface in welds progressively changes as a function of the traveling speed of the heat source whereas it exclusively depends on the time for an ingot. (iv) The movement of the liquid metal in a weld pool is greater than in an ingot due to the Lorentz forces which create turbulence within the molten metal (Grong, 1997). Fig. 13 shows longitudinal views, which depict the direction of solidification of the welds, for a multi-pass welding and MIEA with preheating conditions at 50°C. The arrows indicate the displacing direction of the electric arc.

Fig. 13. Longitudinal top view of the weld metal grain structure at the mid plane for: a) single V groove, b) MIEA preheated at (50°C) (Ambriz, et al, 2010).

The longitudinal macrostructure for the MIEA joint, Fig. 13b, exhibit significant differences with respect to the multi-passes single V groove joint. Irrespective of the preheating condition, the local crystalline growth maintains an angle nearly constant in relation to the moving heat source. The virtual non existence of changes in growth direction means that the local and nominal rates of crystalline growth tend to be equal. This phenomenon yields a significantly different grain structure in the weld metal for the MIEA joint as compared to the structure observed for the single V groove joint. It leads, in fact, to a grain refining effect which is obviously affected by the initial preheating temperature of the joint.

The MIEA joint exhibits signs of heterogeneous nucleation which promotes grain refining. The levels of porosity in the MIEA joints decrease with preheating temperature and are comparatively lower than that obtained in the single V groove joint. Epitaxial solidification is also observed at the fusion line of the MIEA welds, however, competitive columnar growth was restricted instead grain structures alike those observed in the centre of the weld metal were present. The characteristics of solidification observed for the MIEA welds are the result of heterogeneous nucleation which is based on the principle of the formation of a critical radii needed to achieve the energy of formation from potential sites for nucleation such as inclusions, substrates or inoculants (Ti or Zr) (Rao et al., 2008; Ram et al., 2000; Lin et al., 2003). For the MIEA welds, these sites are principally the sidewalls of the joint in conjunction with the content of Ti in the filler and base metal since the significant dilution of base metal favors incorporation of Ti into the weld pool.

#### **4.2 Mechanical properties in aluminum alloys welds 4.2.1 Microhardness**

In order to determine the effect of the welding process in aluminum alloys, a common practice is to perform a microhardness profile in a perpendicular direction to the weld bead, as is showed in Fig. 14. Microhardness measurements give a general idea of the microstructural transformations and the variation of the local mechanical properties (Ambriz et al. 2011) produced after a welding process in aluminum alloys.

preheating conditions at 50°C. The arrows indicate the displacing direction of the electric

Fig. 13. Longitudinal top view of the weld metal grain structure at the mid plane for: a)

The longitudinal macrostructure for the MIEA joint, Fig. 13b, exhibit significant differences with respect to the multi-passes single V groove joint. Irrespective of the preheating condition, the local crystalline growth maintains an angle nearly constant in relation to the moving heat source. The virtual non existence of changes in growth direction means that the local and nominal rates of crystalline growth tend to be equal. This phenomenon yields a significantly different grain structure in the weld metal for the MIEA joint as compared to the structure observed for the single V groove joint. It leads, in fact, to a grain refining effect

The MIEA joint exhibits signs of heterogeneous nucleation which promotes grain refining. The levels of porosity in the MIEA joints decrease with preheating temperature and are comparatively lower than that obtained in the single V groove joint. Epitaxial solidification is also observed at the fusion line of the MIEA welds, however, competitive columnar growth was restricted instead grain structures alike those observed in the centre of the weld metal were present. The characteristics of solidification observed for the MIEA welds are the result of heterogeneous nucleation which is based on the principle of the formation of a critical radii needed to achieve the energy of formation from potential sites for nucleation such as inclusions, substrates or inoculants (Ti or Zr) (Rao et al., 2008; Ram et al., 2000; Lin et al., 2003). For the MIEA welds, these sites are principally the sidewalls of the joint in conjunction with the content of Ti in the filler and base metal since the significant dilution of

In order to determine the effect of the welding process in aluminum alloys, a common practice is to perform a microhardness profile in a perpendicular direction to the weld bead, as is showed in Fig. 14. Microhardness measurements give a general idea of the microstructural transformations and the variation of the local mechanical properties

(Ambriz et al. 2011) produced after a welding process in aluminum alloys.

single V groove, b) MIEA preheated at (50°C) (Ambriz, et al, 2010).

which is obviously affected by the initial preheating temperature of the joint.

base metal favors incorporation of Ti into the weld pool.

**4.2 Mechanical properties in aluminum alloys welds**

**4.2.1 Microhardness**

arc.

Fig. 14. a) Microhardness profile and b) microhardness map in 6061-T6 aluminum alloy welded by MIEA. Note that 1HV=9.810-3 GPa (Ambriz, et al, 2011).

Fig. 14a presents the Vickers hardness number profile in 6061-T6 aluminum alloy welds obtained by MIEA. A significant difference between the hardness number of the weld material and HAZ with respect to the base material is observed. Also, at the limit between the HAZ and the base metal, it is noted the presence of a soft zone which is formed nearly symmetrically in both sides of the welded joints. It should be note that the hardness obtained in this zone represents roughly 57 % of the hardness number of the base material. This seems to indicate that the tensile mechanical properties after welding process will be greatly different. Fig.14b visualizes the location of the soft zone highlighted by the Vickers hardness profile represented in Fig. 14a, by means of a hardness mapping. In this figure, the hardness values for each zone of the welded joint are well-defined. It is clear that in the soft zone (HAZ) the hardness number range is between 0.55 to 0.7 GPa. This soft zone results from the thermodynamic instability of the " needle-shaped precipitates (hard and fine precipitates) promoted by the high temperatures reached during a fusion welding process (Myhr et al., 2004). Indeed the temperatures reached during the welding process are favorable to transform the ' phase, rod-shaped, according to the transformation diagram for the 6061 alloy.

#### **4.2.2 Tensile properties**

The individual mechanical behaviour of the base metal, weld metal, HAZ and welded samples in as welded condition for 6061-T6 aluminum welds by MIEA is shown in Fig. 15 as stress as function of strain graph. From Fig. 15, it can be observed that the experimental results for the base metal are in agreement with nominal values found in the literature for 6061-T6 alloy (American Society for Metals Fatigue and Fracture, 1996). The tensile properties of the sample obtained from the HAZ presents a 41% and a 19 % reduction of the ultimate strength with respect to the base metal and weld metal respectively. The loss of mechanical strength commonly referred to as over-aging, when welding a 6061-T6 alloy is a fairly well understood phenomenon and it is explained in terms of the precipitation sequence (Dutta Allen, 1991). During welding, however, the base metal adjacent to the fusion line is subjected to a gradient of temperature imposed by the welding thermal cycle. At certain distance from the fusion line, the cooling curve crosses the interval of temperatures between 383 to 250 °C in which the *'* phase, rod-shaped, is stable. It is thus the transformation of *''* into *'* the responsible of the decrease in hardening of the matrix due to the incoherence of the *'* phase caused by the thermodynamic instability of *''* in a welding process. It is worthy to mention that the MIEA further increased the mechanical strength of this alloy as compared to the IEA joint (Ambriz et al, 2009).

Fig. 15. True stress-strain curves for 6061-T6 plates, weld metal, HAZ and as welded (Ambriz et al, 2010).

#### **4.2.3 Fatigue crack growth (FCG)**

Fatigue behavior of aluminum alloys welded by conventional process and MIEA has been investigated (Ambriz et al. 2010b; Branza et al., 2009; Seto et al., 2004). Fig. 16a presents the crack length as a function of number of cycles for base metal, weld metal and HAZ in 6061- T6 welds by MIEA, for *P* equal to 2.5 and 3.0 kN. In general, the *a-N* curves showed in Fig. 16a reveal a notable difference in terms of crack length for each material as a function of the number of cycles, nevertheless the small change in *P* (Ambriz et al., 2010b). Experimental results for *a*, were plotted in *da/dN* versus ∆*K* graphs according to Paris law:

$$\frac{da}{dN} = \mathbf{C} \left(\boldsymbol{\Lambda} \mathbf{K}\right)^{\mathrm{II}} \tag{4}$$

where *C* and *n* are constants obtained directly from the fitting curve. Fig. 16b-d, presents the FCG data obtained for the base metal, weld metal and HAZ in MIEA, as well as the comparison with Friction Stir Welding (FSW) data found in the literature (Moreira et al., 2008). Fig. 16b, shows the FCG for base metal in both directions. This graph shows that the microstructure aspect (anisotropy) does not have an important influence in terms of FCG as could be expected, taking into consideration that yield strength in the base metal parallel to rolling direction is higher than transverse direction. However this is not the case for the weld metal and HAZ (Figs. 16c-d), in which the crack tend to propagate faster than base metal. Under this scenario, the FCG behavior for base metal (L-T) was taken as a basis to perform a comparative table between the weld metal and HAZ of MIEA and FSW. Table 2,

welding process. It is worthy to mention that the MIEA further increased the mechanical

Fig. 15. True stress-strain curves for 6061-T6 plates, weld metal, HAZ and as welded

results for *a*, were plotted in *da/dN* versus ∆*K* graphs according to Paris law:

*dN*

Fatigue behavior of aluminum alloys welded by conventional process and MIEA has been investigated (Ambriz et al. 2010b; Branza et al., 2009; Seto et al., 2004). Fig. 16a presents the crack length as a function of number of cycles for base metal, weld metal and HAZ in 6061- T6 welds by MIEA, for *P* equal to 2.5 and 3.0 kN. In general, the *a-N* curves showed in Fig. 16a reveal a notable difference in terms of crack length for each material as a function of the number of cycles, nevertheless the small change in *P* (Ambriz et al., 2010b). Experimental

*da <sup>n</sup> C K*

where *C* and *n* are constants obtained directly from the fitting curve. Fig. 16b-d, presents the FCG data obtained for the base metal, weld metal and HAZ in MIEA, as well as the comparison with Friction Stir Welding (FSW) data found in the literature (Moreira et al., 2008). Fig. 16b, shows the FCG for base metal in both directions. This graph shows that the microstructure aspect (anisotropy) does not have an important influence in terms of FCG as could be expected, taking into consideration that yield strength in the base metal parallel to rolling direction is higher than transverse direction. However this is not the case for the weld metal and HAZ (Figs. 16c-d), in which the crack tend to propagate faster than base metal. Under this scenario, the FCG behavior for base metal (L-T) was taken as a basis to perform a comparative table between the weld metal and HAZ of MIEA and FSW. Table 2,

(4)

*'* the responsible of the decrease in hardening of the

*'* phase caused by the thermodynamic instability of

matrix

> *''* in a

the transformation of

(Ambriz et al, 2010).

**4.2.3 Fatigue crack growth (FCG)**

due to the incoherence of the

*''* into

strength of this alloy as compared to the IEA joint (Ambriz et al, 2009).

presents the crack growth rate, *da/dN* and the stress intensity factor *K*, for base metal, weld metal and HAZ corresponding to a critical crack length in MIEA welds. For comparison effects, values for *da/dN* in MIEA were taken to compute the *K* in FSW.

The results presented in Table 2 indicate that, there is an important difference in *K* for weld metal and HAZ, independently of the welding process. In this way, it should note that *K* for weld metal in MIEA represents only 57% of the base metal, unlike the *K* for weld metal in FSW, which reach a 79% with respect to base metal. This means that FCG rate are higher in MIEA weld metal than FSW, as can be seen in Fig. 16c. This behavior is totally related to the joining processes; it means that MIEA is a welding technique based on a fusion welding process that employs a high silicon content filler metal, which produces a self grain refining, but a brittle microstructure in the weld metal (Ambriz et al., 2010b). On the other hand, FSW is a solid-state joining process that does not use a filler metal (Nandan et al., 2008). Thus, chemical composition in weld metal is similar to the base metal and microstructural characteristics related to dynamic recrystalization tends to be better than MIEA. In contrast, Fig. 16d shows that FCG rate in MIEA and FSW is similar in the HAZ. The stress intensity factor relation was 64% with respect to base metal. It is noted that thermal effect produced by the microstructural transformation of very fine precipitates needle shape '', to coarse bar shape ' precipitates, has a profound impact in the HAZ crack growth rate. It confirms that, independently of the welding process (MIEA and FSW), the crack growth conditions are directly influenced by the temperature within the HAZ, which is normally above of the aging temperature of the alloy, causing a hardening lost and important decrease in mechanical properties.

Fig. 16. a) crack length as function of number cycles, load ratio R=0.1, b-d) Fatigue crack growth rate as function of stress intensity factor range (Ambriz et al, 2010).

Table 2. Comparison between MIEA and FSW based on a critical crack length. BM = Base metal, *i* corresponds to weld metal or *HAZ for MIEA or FSW* (Ambriz et al, 2010).

#### **5. Conclusions**

The IEA technique and its evolution into the MIEA have emerged as an attractive alternative to weld a number of materials with peculiar microstructural characteristics that have a positive impact in the mechanical and stress corrosion cracking (SCC) behaviors. Between these material are aluminum metal matrix composites (MMCs) reinforced with TiC, SiC and Al2O3 particles and monolithic materials such as carbon steels (API X-60 and X-65), aluminum and its alloys, such as 6061, 2014 and 359.

With the IEA welding process, a columnar microstructure with a fine grain, more homogeneous structure and a small HAZ is obtained. In this process, the heat transfer developed does not affect the base metal as much as the direct electric arc process does, but the heat input is enough to partially melt the base metal yielding a weld profile with a high depth to-width ratio. In addition to the macroscopic features obtained, the use of this technique might lead to obtaining a fine microstructure in the weld for metallic materials and improved mechanical properties. The evolution into the MIEA has marked a significant progress regarding mechanical behavior as a result of the reduced thermal affection in heat treatable aluminum alloys.

#### **6. References**

Base metal MIEA FSW

metal, *i* corresponds to weld metal or *HAZ for MIEA or FSW* (Ambriz et al, 2010).

aluminum and its alloys, such as 6061, 2014 and 359.

(mm/cycle) 1.981×10-3 1.0×10-3 1.413×10-3 1.0×10-3 1.413×10-3 *K* (MPa m1/2) 30.41 17.27 19.46 23.98 19.51 (*da/dN)i/(da/dN)BM* 1.0 0.50 0.71 0.50 0.71 (∆K)i/(∆K)BM 1.0 0.57 0.64 0.79 0.64 Table 2. Comparison between MIEA and FSW based on a critical crack length. BM = Base

The IEA technique and its evolution into the MIEA have emerged as an attractive alternative to weld a number of materials with peculiar microstructural characteristics that have a positive impact in the mechanical and stress corrosion cracking (SCC) behaviors. Between these material are aluminum metal matrix composites (MMCs) reinforced with TiC, SiC and Al2O3 particles and monolithic materials such as carbon steels (API X-60 and X-65),

With the IEA welding process, a columnar microstructure with a fine grain, more homogeneous structure and a small HAZ is obtained. In this process, the heat transfer developed does not affect the base metal as much as the direct electric arc process does, but the heat input is enough to partially melt the base metal yielding a weld profile with a high depth to-width ratio. In addition to the macroscopic features obtained, the use of this technique might lead to obtaining a fine microstructure in the weld for metallic materials and improved mechanical properties. The evolution into the MIEA has marked a significant progress regarding mechanical behavior as a result of the reduced thermal affection in heat

Ahearn, J.; Cook, C. & Fishman, S. (1982). Fusion welding of SiC-reinforced Al composites. *Metal construction*, Vol. 14, Issue 4 (April 1982), pp. 192-197, ISSN 03077896 Ambriz, R.; Barrera, G. García, R. (2006). Aluminum 6061-T6 welding by means of the

Ambriz, R.; Barrera, G.; García, R. López V.H. (2008). Microstructure and heat treatment

Ambriz, R.; Chicot, D.; Benseddiq, N.; Mesmacque, G. de la Torre, S. (2011). Local

Ambriz, R.; Mesmacque, G.; Ruiz, A.; Amrouche, A. & López, V.H. (2010). Effect of the

(May-Jun 2011), pp. 307-315, ISSN 0997-7538

modified indirect electric arc process. *Soldagem and Inspecao*, Vol. 11, Issue 1, pp.

response of 2014-T6 GMAW welds obtained with a novel modified indirect electric arc joint. *Soldagem and Inspecao*, Vol. 13, Issue 3, (Jul-sep 2008), pp. 255-263, ISSN

mechanical properties of the 6061-T6 aluminium weld using micro-traction and instrumented indentation. *European Journal of Mechanics A/Solids*, Vol. 30, Issue 3,

welding profile generated by the modified indirect electric arc technique on the

*da dN* /

**5. Conclusions**

treatable aluminum alloys.

10-17

0104-9224

**6. References**

Weld metal HAZ Weld metal HAZ

fatigue behavior of 6061-T6 aluminum alloy. *Materials Science and Engineering A*, Vol. 527, Issue 7-8, (Mar 2010), pp. 2057-2064, ISSN 0921-5093

Garcia, R.; Lopez, V. H.; Bedolla, E.; Manzano, A. (2007). Welding of aluminium by the MIG

García, R.; López, V.H.; Bedolla E. & Manzano A. (2002). MIG welding process with indirect

García, R.; López, V.H.; Bedolla E. & Manzano A. (2003). A comparative study of the MIG

*Journal of Materials Science*, Vol. 38, (2003), pp. 2771– 2779, ISSN 0022–2461 García, R.; Manzano, A.; López, V.H. & Bedolla, E. (2002). Comparative welding study of

Garcia, R; López, V.H.; Kennedy, A. & Arias, G. (2007). Welding of Al-359/20%SiCP metal

Gittos, N.; & Scott M. (1981). Heat-affected zone cracking of Al-Mg-Si alloys. *Welding Journal*,

Grong, O. (1997). *Metallurgical Modeling of Welding*, (Second Edition), Maney Materials

Guiterrez, L.; Neye, G. & Zschech, E. (1996). Microstructure, hardness profile and tensile

Heinz, A.; Haszler, A.; Keidel, C.; Moldenhauer, S.; Benedictus, R. & Miller, W. (2000).

Kerr, H. & Katoh, M. (1987). Investigation of heat-affected zone cracking of GMA welds of

Kluken, A. & Bjorneklett, B. (1997). A study of mechanical properties for aluminum GMA weldments. *Welding Journal*, Vol. 76, No.2, pp. 39-44, ISSN 0043-2296 Lienert, T.; Brandon, E. & Lippold, J. (1993). Laser and Electron beam welding of a SiCP

Liu, G.; Murr, L.; Niou, C.; McClure, J. & Vega, F. (1997). Microstructural aspects friction-stir

Liu, W. J.; Tian, X. & Zhang, X. (1996). Preventing weld hot cracking by synchronous rolling during weld. *Welding Journal*, Vol.75, No.9, pp. 297-s-304-s, ISSN 0043-2296

*materialia,* Vol. 28, Issue 11 (Jun 1993), pp. 1341-1346, ISSN 0956-716X Lin, D.; Wang, G. Srivatsan T. (2003). A mechanism for the formation of equiaxed grains

*Science and Engineering A*, Vol. A280, No.1, pp. 102-107, ISSN 0921-5093 Huang, C. & Kou, S. (2004). Liquation cracking in full-penetration Al-Mg-Si welds. *Welding*

*of Materials Science,* Vol. 42, pp. 7794-7800, ISBN 0022-2461

Vol.49, No.2, pp. 96-s-103-s, ISSN 0043-2296

*Journal*. Vol. 83, No. 4, pp. 111s-122s, ISSN 0043-2296

351, Issue 1-2, (Jun 2003), pp. 304-309, ISSN 0921-5093

Science, ISBN<PHONE_NUMBER>, London.

7956-7963, (Jun 2007), ISSN 0022-2461

0261–8028

932 a 937, ISSN 1073-5615

121s, ISSN 0043-2296

259s, ISSN 0043-2296

1359-6462

process with indirect electric arc (MIG-IEA), *Journal of Materials Science*, Vol.42, pp.

electric arc. *Journal of Materials Science Letters*, Vol. 21, (2002), pp. 1965-1967, ISSN

welding of Al/TiC composites using direct and indirect electric arc processes.

metal matrix composites with the MIG welding process, using direct and indirect electric arc. *Metallurgical and Materials transactions B*, Vol. 33B, (December 2002), pp.

matrix composites by the novel MIG process with indirect electric arc (IEA). *Journal*

strength in welds of AA6013 T6 extrusions. *Welding Journal*, Vol.75, No.4, pp. 116s-

Recent development in aluminium alloys for aerospace applications. *Materials*

Al-Mg-Si alloys using the varestreint test. *Welding Journal*, Vol.66, No. 9, pp. 251s-

reinforced aluminum A-356 metal matrix composite, *Scripta metallurgical et*

in welds of aluminum-lithium alloy 2090. *Materials Science and Engineering A*, Vol.

welding of 6061-T6 aluminum. *Scripta Materialia*, Vol.37, No.3, pp. 355-361, ISSN

## **Part 2**

**Arc Welding Automation**

44 Arc Welding

Urena A.; Rodrigo P.; Gil L.; Escalera M. & Baldonedo J. (2001). Interfacial reactions in an Al-

Urena, A.; Rodrigo, P.; Gil, L.; Escalera, M. & Baldonedo, J. (2000). Interfacial reaction in Al-

*science*, Vol.32, pp. 419-428, ISBN 0022-2461

Cu-Mg (2009)/SiCw composite during liquid processing - Part II - Arc welding. *Journal of materials science*, Vol. 36, Issue 2 (Jan 2001), pp. 429-439, ISSN: 0022-2461

Cu-Mg (2009)/SiCP composite during liquid processing Part 1, *Journal of materials*

### **Arc Welding Automation**

Eduardo José Lima II and Alexandre Queiroz Bracarense *Universidade Federal de Minas Gerais Brazil*

#### **1. Introduction**

It is very well known that the use of robotics in the industry increases productivity and quality in many aspects. It is also well known that some adjustments have to be made to grant payback for the investment and to reach the expected results. Today, the robotization is known as an alternate technique for production, increasing qualitative and quantitative competence of its industries. One point that has been observed in some applications is that inadequate procedure for robot application in welding process reflects in insatisfactory results. These procedures are excessive time expending for implementation, material loss, reworking and poor weld quality.

The main problem detected is that the number of experts in welding "and" robotics is still reduced and some industries are investing in robot without any planning or orientation, believing that the robot will solve all their problems. The results have been disastrous. In many cases even though the weld appearance is acceptable and the welding time is significantly reduced, the weld quality is poor. The experts in robotics know very well what they are doing. However, many of them do not have experience in welding to understand that many features related to welding physics and metallurgy have to be considered when a welding procedure has to be implemented. Because of this, some small and medium industries with large potential for robotization are holding investment and/or postponing it until its adaptation to implement the robot in their production line. It is believed however that very soon many small industries will have their own robot.

#### **2. The paradigm of welding automation**

Since the beginning, welding is a process that depends on the welder skills. This relation is so direct that the classification, according to the application methods, is based on the degree of control of the activities related to welding that depends on the human interference. These application methods are classified as manual, semi-automatic, mechanized, automatic, robotic and with adaptive control, according to American Welding Society (AWS).

This classification can be better understood when an agent is established (Table 1) to execute the normal activities involved to realize arc welding (Cary, 1994).

The manual welding is defined, according to the American Welding Society (AWS, 2001) as "welding where the torch or the electrode holder are carried or manipulated by human hands". In other words when the tasks, related with the execution and continuous control of the welding, are made by the hands of the human and are under responsibility of him.

Table 1. Application methods of welding process, adapted from (Welding Handbook, 2001).

Semi-automatic welding is defined as "manual welding with equipment that automatically control one or more welding conditions". The welder manipulates the torch while the wire/electrode is automatically fed by the machine.

Mechanized welding is defined as "welding with equipment that requires manual adjustments in response to visual observation during welding, with the torch or electrode holder carried by a mechanical system". The welder participation in this process consist in adjusting the parameters as he observe the operation.

Automatic welding is defined as "welding with equipment that only requires occasional observation and/or no observation of the weld and no manual adjustments". The function of the welder is only turn the machine on to begin the welding cycle and occasionally check the procedure.

Robotic welding is defined as "welding that is executed and controlled by a robotic equipment". In robotic welding and automatic welding the function of the welder is to guarantee the weld quality by performing periodic inspections of the results identifying weld discontinuities. When those are find, maintenance and programming must be done to fix such problems.

Welding with adaptive control is defined as "welding with equipment that has a control system that automatically determine changes in welding conditions and act under the equipment with appropriated action to do adjustments". In this process, sensors are used to detect problems and the controller performs the necessary changes in welding parameters,

maintenance Human Machine Machine Machine Machine Machine

Wire feeding Human Machine Machine Machine Machine Machine

Human Human Human Machine Machine

Table 1. Application methods of welding process, adapted from (Welding Handbook, 2001). Semi-automatic welding is defined as "manual welding with equipment that automatically control one or more welding conditions". The welder manipulates the torch while the

Mechanized welding is defined as "welding with equipment that requires manual adjustments in response to visual observation during welding, with the torch or electrode holder carried by a mechanical system". The welder participation in this process consist in

Automatic welding is defined as "welding with equipment that only requires occasional observation and/or no observation of the weld and no manual adjustments". The function of the welder is only turn the machine on to begin the welding cycle and occasionally check

Robotic welding is defined as "welding that is executed and controlled by a robotic equipment". In robotic welding and automatic welding the function of the welder is to guarantee the weld quality by performing periodic inspections of the results identifying weld discontinuities. When those are find, maintenance and programming must be done to

Welding with adaptive control is defined as "welding with equipment that has a control system that automatically determine changes in welding conditions and act under the equipment with appropriated action to do adjustments". In this process, sensors are used to detect problems and the controller performs the necessary changes in welding parameters,

along the joint Human Human Machine Machine Machine

Human Human Human Do not

Human Human Machine Machine Machine Machine

Machine (with pre programed track)

occours

automatic Mechanized Automatic Robotic Adaptive

control

(with sensor)

(with sensor)

(with sensor)

Machine (with sensor)

Machine (with sensor)

Machine (with sensor)

Machine (with sensor)

(robot)

Machine (robot)

(robot)

Do not occours

Manual Semi-

along the joint Human Human Human

wire/electrode is automatically fed by the machine.

adjusting the parameters as he observe the operation.

Activities

Arc start and

Heat control to obtain penetration

Arc motion

Guide the arc

manipulation to direct the arc

Arc corrections to compense erros

the procedure.

fix such problems.

Torch

in real time, to produce sound welds. This type of welding is performed without intervention and supervision of the human.

Accord to the classification presented, the stick electrode welding is a manual process since the welder is responsible for all the activities, while GMAW is a semi-automatic process. This is so because the arc start and maintenance and wire feeding is made by the machine while the torch manipulation is made by the welder. When this manipulation is made by a mechanical device the processes is classified as mechanized.

Independent of the automation degree, its focus is costs reduction by reducing the number of people involved in the production and increasing productivity and quality of the final product by the rational control of the process parameters. An automatic equipment may be projected and programmed to perform a unique task (fix automation) or may be projected for multitasks, by programming, allowing to perform distinct tasks accord with the manufactured product (flexible automation) (Welding Handbook, 2001; Romano, 2002).

In manufacturing area, the term "automation" means that all the functions or steps of an operation are executed or controlled, in sequence, by mechanical and/or electronics devices replacing the human efforts, observation and decision.

The automation involves more than equipments or control by computer and may or may not include charge or discharge of components in an operation. The automation may be partial, with some functions or steps executed manually or may be total, where all the functions or steps are executed by the equipment, in a certain sequence without any adjustment by the operator (Romano, 2002).

To properly classify the automation level of a given process, the first step is to define the activities related to it. The responsible for the execution (execution agent), for the control (control agent) and for the sequence of activities (sequential agent) must be defined. Additionally, it is necessary to define which activities need to be treated as isolated and which must be included in the process operational cycle.

In the arc welding processes, the first activity to execute a weld is to specify which welding procedure must be used. This definition is also known as WPS – Welding Procedure Specification. Considering used material and the weld morphology wanted, the welding parameters related to the process may then be determined. The executor, controller and sequence specification agent of a WPS is the human. Even with a help of a computational program to choose the best parameters or the ones that will be used for the self adjustment of the machine, the final decision for the WPS is the human decision. Between the WPS elaboration and the welding beginning, a time interval for procedure preparation is necessary. Nowadays, this preparation occurs independent of the application method to be used and human direct interference are often necessary.

There is still a lack of a good system to automatically prepare the WPS and, from its preparation, immediately initiate the welding. Of course this system must be universal, since among the parameters to be chosen for the WPS there are the welding processes. It is, however, expected that such activities (WPS and welding start) be treated as isolated activities, because of the human involvement on them. The activities related to the welding cycle, the ones that will define the degree of automation of the process, must be the ones that allows the instantaneous sequence of the process.