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Effect of the weld thermal cycles of the modifiedindirect electric arc on the mechanical properties ofthe AA6061-T6 alloyRicardo R. Ambriz a , Gerardo Barrera a , Rafael García a & Victor H. López aa Institute of Metallurgical Investigations , Universidad Michoacana de San Nicolás deHidalgo , Morelia, Mich., MexicoPublished online: 12 Feb 2010.

To cite this article: Ricardo R. Ambriz , Gerardo Barrera , Rafael García & Victor H. López (2010) Effect of the weld thermalcycles of the modified indirect electric arc on the mechanical properties of the AA6061-T6 alloy, Welding International, 24:4,321-328, DOI: 10.1080/09507110903568778

To link to this article: http://dx.doi.org/10.1080/09507110903568778

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Effect of the weld thermal cycles of the modified indirect electric arc on the mechanicalproperties of the AA6061-T6 alloy

Ricardo R. Ambriz1, Gerardo Barrera2, Rafael Garcıa3 and Victor H. Lopez

Institute of Metallurgical Investigations, Universidad Michoacana de San Nicolas de Hidalgo, Morelia, Mich., Mexico

(Received 30 January 2008; final version received 6 August 2008)

Results of temperature measurements during welding of 12.7 mm thick AA6061-T6 alloy plates by modified indirect electricarc (MIEA) are presented. This study describes the thermal cycles in the heat-affected zone (HAZ) and also in the fusionzone. Depending upon the position of the transducers, the maximum temperatures measured in the HAZ ranged from 308 to6938C, these measurements were related with the tensile test results and the failure zone reported previously by the authors(Ambriz et al., S&I 2006;11:10–17). It was observed that there is a decrease in the mechanical strength of the welded joints,due to the microstructural changes undergone by the AA6061-T6 alloy in which formation of the b 0 occurs according to thetime–temperature transformation diagram. The inherent cooling conditions of the weld pool observed for the MIEAtechnique (single welding pass) have made it possible to establish the characteristics of solidification and microstructure fora specific cooling rate.

Keywords: aluminium alloy; melted pool; heat-affected zone; transformations; overageing; weld thermal cycle

1. Introduction

The measurement of temperature is a very important aspect

of processes in which the thermal effects cause changes

that affect the behaviour of the materials, as is the case with

welding, where the heat supplied due to the electric arc

produces significant microstructural changes on the

welded materials and, consequently, on their mechanical

properties2,3. There are a wide variety of sensors for

measuring temperature, which are classified according to

the range of measurement and type of application.

Thermocouples are the temperature measurement instru-

ments par excellence; they are thermoelectric elements that

convert a change in temperature into a change of voltage

(the Seebeck effect); they are manufactured from various

materials, such as Chromel–Alumel (type K), which

operate in a range (where behaviour is linear) from 2270

to 12608C4. The measurement of temperature during the

welding process of heat-treatable aluminium alloys is of

great interest, due to the sensitivity of these alloys to

temperature changes that give rise to a loss of their

hardening states, with the consequent reduction in their

mechanical properties5. Another important point is

liquefaction cracking, found within a high-temperature

region in the heat-affected zone (HAZ)6. Various authors

have studied the effect of temperature on the mechanical

properties and the microstructural changes in heat-

treatable aluminium alloys3,7,8.

The recording of temperature variations during

the welding process, in the molten zone, the HAZ and

the base material, has the objective of establishing the

microstructural changes that the alloy may undergo and

that, inevitably, will be reflected in the mechanical

properties of the welded joint. It should be recalled that

the principal hardening mechanism in heat-treatable

aluminium alloys is the metastable precipitation of

hardening phases, which comprise three stages: solubil-

ization, tempering and ageing (artificial or natural)9.

The type of solidification of fusion welding is

completely related to the heat supply, the chemical

composition of the welding metal, the rapidity of crystal-

line growth, the welding speed and the profile of the molten

weld pool in such a way that the grain size of the molten

base metal at the fusion limit acts like a substrate for the

growth of columnar grains10. Additionally, the direction of

growth of the columnar grains changes continually from the

fusion line towards the centre of the weld, due to the

corresponding change in the direction of the maximum

temperature gradient in the molten pool. This is a very

particular problem for high-energy welding processes such

as submerged arc welding and gas metal arc welding

(GMAW), where the growth of the grain of the base metal is

considerable. Even more so during multipass welds, where

the columnar grains may re-nucleate within their limits

from one welding pass to the next.

Recently, the authors have experimented with joint

design in indirect electric arc (IEA) welding11 – 15 in order

to weld materials composed of metal matrices and other

monolithic materials. The technique consists of position-

ing small feed sheets along the joint of squared edges. The

electric arc is established indirectly over the plates,

forming a liquid pool of welding that, due to gravity and

the impulsion of the electric arc, feeds down towards the

ISSN 0950-7116 print/ISSN 1754-2138 online

q 2010 Taylor & Francis

DOI: 10.1080/09507110903568778

http://www.informaworld.com

Welding International

Vol. 24, No. 4, April 2010, 321–328

Selected from Revista de Metalurgia 2009 45(1) 42–51

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bottom of the joint, making it possible to weld thicknesses

of 12.7 mm in a single welding pass. This technique,

however, has the disadvantage that the remaining feed

sheets have to be removed. Under the same schema, the

authors modified this joint. Instead of using the feed

sheets, they machined a small flap in the upper part of the

sheets to be welded, so that this design came to be called

modified indirect electric arc (MIEA) welding.

The present investigation was carried out in order to

determine the relationship between the temperature

measurements and the effect on the mechanical resistance

of the joints, due to the microstructural changes due to the

transformation of intermetallics of the Mg2Si-type using

MIEA welding. In the same context, the conditions and

characteristics of the cooling of the pool of liquid metal

were studied, which are also influential in the mechanical

behaviour of a welded joint. It is important to emphasize

that for thickness of 12.7 mm, it is necessary to apply

several welding passes, which leads to a greater heat supply

to the base material and, consequently, a greater degree of

heat effect and softening in the HAZ of a precipitation-

hardened aluminium alloy. Thus, it proves beneficial to use

a joint design that allows the joining of these alloys in a

single pass, while reducing the loss of hardening. Also, it is

important to note that the mechanization of the new joint is

no more complicated or costly than the preparation of a

simple V joint.

2. Experimental development

2.1 Operational variables, welding conditions andmechanical properties

The base metal used was an aluminium alloy 6061-T6

(Al–Si–Mg) in sheets, with a thickness of 12.7 mm. In

order to evaluate the mechanical properties of the welded

joints, sheets of 70 mm width and 150 mm length were used

and traction specimens were mechanized in accordance

with ASTM standard B557M-94. The results of these tests

were presented in a work cited above1. With reference to

the zones where the failure of the traction specimens took

place and the microhardness measurements made in the

HAZ, the authors decided to carry out the measurement of

the thermal cycle welding sheets with a thickness of 25 mm

each. Moreover, this width made it possible to place the

thermocouples with precision. The filler material was a

commercial electrode with a high silicon content

(ER4043) and a diameter of 1.2 mm. Table 1 shows the

chemical composition of the materials used, which was

obtained using atomic absorption spectroscopy.

A semiautomatic GMAW process was used, with

100% argon as the shielding gas, with a flow of 23.6 l/min.

The operative variables were adjusted in order to obtain a

metal transference via pulverization, with an approximate

current of 230 A, a separation of the contact tip and the

base material of 20 mm, inverse polarity (CDEP), constant

voltage of 23 V and a displacement speed of 3.6 mm/s of

the heat source; in this case, a joint preparation of the flap

style, which was named MIEA, using a preheat of 50, 100

and 1508C1. The joint design and the dimensions thereof

are shown in Figure 1.

The base material, filler material and welded joints

were traction tested (using at least three specimens)

according to the recommendations of ASTM standard B

557M-9416. Table 2 shows the results obtained for the base

material and filler material, while those corresponding to

the welds are presented in the analysis and discussion of the

results. Additionally, microhardness measurements were

carried out on the base material and welded joints, applying

a load of 0.1 N for 15 s. The average microhardness value

for the base material was 152.5HV0.01, while the welded

joints were evaluated using microhardness scans and

correlated with their welding profile.

Table 1. Chemical composition of the 6061 alloy and ER4043 filler wire, wt.%.

Alloy Si Fe Cu Mn Mg Cr Zn Ti Al

6061-T6 0.561 0.289 0.310 0.052 0.986 0.067 0.024 0.018 Bal.ER4043 5.25 0.8 0.30 0.05 0.05 – 0.10 0.20 Bal.

Figure 1. Joint design and thermocouple location for measuringthe temperature in the liquid metal pool.

Table 2. Mechanical properties of the base material and fillerwire.

AlloyFlow strength

(MPa)Stress resistance

(MPa)Lengthening

(%)

6061-T6 300 328 14ER4043 164 190 –

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2.2 Temperature measurement in the HAZ

In order to avoid interference in the heat transference

patterns at the time of completing the weld, small

thermocouples, type-K, with a wire diameter of 0.3 mm,

were set up. Silver covers were manufactured, with an

external and internal diameter of 1.5 and 1 mm,

respectively, with the thermocouples being situated within

the said covers or isolators. The isolation of the

thermocouples was completed by filling the internal

diameter with a silicon oxide (SiO2) ceramic, as shown in

Figure 2.

The placement of the thermocouples for the

measurement of temperature in the HAZ was established

by fixing a system of rectangular coordinates (X, Y, Z), the

origin of which was located in the upper face of the base

metal sheets and in the centre of the joint preparation

(Figure 3). A total of 10 thermocouples were placed in

the alloy sheets to be welded. Table 3 shows the positions

of each of these.

2.3 Temperature measurement in the liquid weld pool

A type-K thermocouple, with a wire diameter of 0.8 mm,

was located at the bottom of the prepared joint, using a

hole in the supporting plate, in order to obtain temperature

measurements in the molten pool, the location of which is

shown in Figure 1. It is well known that the cooling speed

in the fusion zone is much quicker than that of an ingot17.

Consequently, it is necessary to take into consideration a

sampling speed that makes it possible to provide

information regarding the solidification phenomenon, as

a result of which the signals were acquired at 60 readings

per second; i.e. the signals were converted from analog to

digital at 60 Hz.

2.4 Digitalization of the signals

In order to record the signals from the thermocouples, a

card connected with one of the PCI ports of the computer

was used, with 16 different analog inputs. This includes a

thermocouple conditioning unit. For the data acquisition

phase, programs written in the graphic programming

language G, better known as ‘Virtual Instrumentation’

(LabVIEW 8.2), were used, within the Windows XPw

operating system.

3. Results and discussion

3.1 Temperature in the HAZ and effect on mechanicalresistance

Table 4 shows the maximum temperatures measured for

the different preheat conditions. It can be seen that the

temperature intervals increase according to the depth at

which the sensors are found relative to the heat source, as

well as the preheat conditions used. For example, the

maximum temperatures measured for position 1 are 308.8,

450.8 and 534.18C. However, this behaviour is not

complied with in all cases, e.g. for the maximum

temperature measured in position 2 with preheat at

1008C, which saw an increase of 458C over the maximum

temperature measured for a preheat of 1508C. This

phenomenon can be explained according to the disalign-

ment of the joint relative to the heat source and the nature

of the arc itself, which tends to maintain its stability

Figure 2. Sketch representation of a K thermocouple fortemperature measurement in the HAZ.

Figure 3. Coordinate system for the location of thethermocouples.

Table 3. Location of the thermocouples.

Thermocouple X (mm) Y (mm) Z (mm)

1 4.0 10.0 29.72 6.0 42.5 26.73 3.0 75.0 21.24 8.0 107.5 210.75 10.0 140.0 23.76 212.0 10.0 28.77 25.0 42.5 24.78 27.0 75.0 22.79 211.0 107.5 210.710 29.0 140.0 26.7Bottom 0 72.0 29.2

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towards one side, causing it to have greater molten mass

on one of the walls of the joint, giving rise to higher

temperature measurements. Moreover, the effect of heat

conduction due to the fusion and solidification of the base

material and the heat supplied have a very significant

effect on the temperature measurements of the sensors

found at the end of the sheets or at different heights

respective to the heat source18.

In Figure 4, the curves corresponding to temperature

measurements within the HAZ are shown for position 6,

corresponding to 12-mm distance from the central axis of

the joint preparation, in accordance with Table 3, for 500,

100 and 1508C of preheat. It can be seen that the maximum

temperature peaks reach a higher value, as the preheat is

increased. The values of the heating and cooling rates

calculated are shown in Table 5.

It should be noted that the behaviour of the cooling rate

for each of the curves after a certain time has passed

(approximately 200 s), and once it has passed the time–

temperature transformation (TTT) curve is practically

identical. This behaviour offers sufficient information to

explain why the traction resistance or maximum strain of

the welded joints appears to be practically the same, as can

be seen in Table 6.

Here, too, the correlation of the thermal cooling cycles

with the TTT curve of the alloy can be seen, corresponding

to the thermocouple furthest from the fusion zone, it being

found that for any preheat condition, the corresponding

cooling cycles fall back to the TTT curve for the formation

of the b 0 phase (Mg2Si). In addition, it is to be expected

that the growth of overaged precipitates (b 0 phase) will be

greater for the condition of 1508C preheat, promoting the

occurrence of the failure zone further from the fusion line

(Figure 5)1.

It is clear that the maximum temperatures reached (in

relation to thermocouple 6) do not exceed the solubil-

ization temperature of alloy 6061, which is found at

around 5308C. However, these temperatures are above the

temperature for the formation of Guinier–Preston zones,

in accordance with the precipitation sequence for alloys of

Al–Si–Mg20. The maximum temperatures reached are

383, 421 and 4508C for their respective preheats, which are

very close to the corresponding values reported by Myhr

et al.7, who established that the transformation of b 00

precipitates to b 0 occurs when the peak temperature

exceeds 3258C, with a time interval of at least 10 s

(a condition fulfilled in the case of the curves of

Figure 4) and at 3908C, the b 0 phase begins to be the

dominant microstructural constituent. The above also

corresponds to the investigations carried out by Malin2

into the relationship that exists between the maximum

temperature measured in the HAZ and the location of

failure in aluminium alloys 6061-T6 extruded and welded

in the conventional manner by means of a simple V joint

preparation, finding that the minimum microhardness

value and fracture location are associated with temperature

peaks at around 3808C.

Table 4. Maximum temperatures measured in the HAZ.

Preheat

Position 508C 1008C 1508C

1 308.8 450.8 534.12 368.7 542.9 497.93 409.5 430.3 611.84 469.8 462.9 577.65 530.7 479.5 553.16 383.1 421.3 450.27 461.7 384.1 577.08 445.9 460.0 574.99 Open 456.4 614.410 610.5 552.2 693.7

Table 5. Heating and cooling rates.

508C 1008C 1508C

Heating (8C/s) 16.8 (1.0088C/min)16.5

(9908C/min)8.0

(4808C/min)

Cooling (8C/s) 0.6 (368C/min)0.8

(488C/min)1.0

(608C/min)

Figure 4. Cooling thermal cycles at 12 mm from the heat sourceand its relation with the TTT curve5,19 for 6061 alloy.

Table 6. Mechanical properties of the welded joints1.

Preheat(8C)

Lengthening(%)

Elasticlimit (MPa)

Tractionresistance (MPa)

Failurezone

50 13.8 102.7 183.0 HAZ100 15.4 101.0 181.7 HAZ150 17.2 106.5 179.3 HAZ

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On the other hand, the b 00 phase with acicular

morphology is that which produces the greatest effect on

the hardening of the alloy 6061-T6, according to that

established by Dutta and Allen20, while the b 0 phase, which

has a bar shape, will grow depending on the temperature

increase in the HAZ giving rise to overageing and,

consequently, its incoherence with the aluminium matrix.

The above is shown in the image of Figure 6, corresponding

to the fracture of one of the MIEA joints, in which a Mg2Si

particle of large size can be seen, among others.

It is clear that the fracture was caused by the softening

of the HAZ by the formation of particles of b 0 and/or b

(equilibrium) precipitates, a product of the thermodynamic

instability of the b 00 precipitates during a fusion welding

process. However, the mechanical properties obtained

after the MIEA joining process (with a single welding

pass) are considerably greater relative to welds carried

Figure 5. Relationship between the microhardness profiles and the failure zone (a) MIEA 508C, (b) MIEA 1008C and (c) MIEA 1508C.

Figure 6. Fracture of the 6061-T6 weld showing Mg2Siparticles.

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out in the conventional manner, with a simple V joint

preparation and three or four welding passes1,21,22.

3.2 Temperature in the molten weld pool

Figure 7 shows the graph of the data obtained using the

data acquisition system for the measurement of tempera-

ture in the molten weld pool, in accordance with the

position of the thermocouple in Figure 1, in which it can be

seen that the process of data acquisition began when the

minimum temperature was 978C (time 0 s) and reached a

maximum temperature of 11578C, corresponding to 1198

readings (19.96 s), as a result of which a temperature

increase of 10608C was obtained.

Carrying out an analysis of the data obtained, it was

observed that, for a total of 615 readings (10.25 s), i.e.

when T ¼ 1118C, the increase in the heating rate was

insignificant, at around 1.388C/s. This increment corre-

sponds to the initial straight line portion of the curve of

Figure 7, as a result of which it is possible to take the time

of beginning temperature measurement at the value of 615

readings, an aspect that becomes obvious due to the fact

that the slope of the curve from this value onwards begins

to be very long. Based on the above, it can be determined

that the time required to reach the maximum temperature

is 9.7 s, with the heating rate being able to be calculated

through the slope of the curve of Figure 7, by means of the

following ratio:

Gt ¼Tp 2 T i

t¼

ð1157 2 111Þ8C

9:7 s¼ 107:88C=s; ð1Þ

where Gt is the heating or cooling rate, Tp is the maximum

temperature and Ti is the initial temperature recorded by

the thermocouple.

The actual distance of the thermocouple relative to the

beginning of the thermal cycle of the weld (beginning of

the arc) was 72 mm (Table 3), which should in theory

correspond to the time at which the temperature is

maximum, i.e. 19.96 s. The above may be corroborated

according to the weld speed used for the welding process

(3.6 mm/s), and by making use of the definition to

determine the speed at that moment, it is possible to

calculate the distance in the following manner:

d ¼ vt ¼ ð3:6 mm=sÞð19:96 sÞ ¼ 71:85 mm: ð2Þ

The value calculated results in a distance that

corresponds, approximately, to the distance at which the

thermocouple was located. Consequently, there is

sufficient evidence to determine that the filler material is

forced, by the effect of gravity, towards the joint of the

plates to be joined and not by means of a possible flow of

molten metal ahead of the heat source (i.e. the heat source

is perpendicular to the temperature sensor). An important

point to take into consideration is that the maximum

temperature measured in the welding metal is the

temperature corresponding to a distance of 3.5 mm

above the support plate and not the temperature at which

the drops of filler material are detached by the welding

process8.

Figure 8 is the graph corresponding to the time interval

in which the temperature is maximum in the liquid weld

metal. Carrying out an analysis according to the fusion

temperature of the filler material (6308C) and its eutectic

temperature (5818C), it is possible to determine the

cooling time for a temperature drop of 498C, correspond-

ing to 0.76 s, as a result of which the cooling rate will be

Gt ¼498C

0:76 s¼ 64:478C=s:

This cooling speed is lower than the high values for a

GMAW process23. However, the degree of previous sub-

cooling experienced in the solidification process promotes

Figure 8. Temperature profile of the weld pool in which thetemperature is maximum and the liquid–solid transformationtakes place.

Figure 7. Profile of the temperature measured in the weld pool.

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the development of a fine grain size, as shown in the

micrograph of Figure 9, corresponding to the central

portion of the weld for an MIEA joint preheated to 1508C.

Once the maximum temperature has been reached, this

begins to decrease due to the movement of the heat source

according to the weld speed of the welding process,

according to what has been established in investigations into

moving heat sources by Rosenthal18. During cooling, energy

begins to be supplied to the lateral walls of the joint

preparation, experiencing a drop in temperature until a

minimum, corresponding to 7378C, is reached, which is

above the fusion point of the base material (6528C) and of the

filler material (6308C). At this point, the lateral walls are

completely fused and consequently a liberation of the latent

heat of fusion occurs, which represents the energy generated

during the liquid–solid transformation, which is absorbed by

the liquid metal, increasing its temperature by around 758C in

a short time, 1.41 s, reaching a maximum of 8128C.

The classical theory for heterogeneous nucleation

explains that the form in which the latent heat of fusion is

dissipated determines the growth mechanism and final

structure of the solidified material. This theory was the basis

for the development of the ideas of Garcia et al.13, who

explained, schematically, the behaviour of the cooling rate in

traditional welding processes and by IEA (Figure 10).

It can be seen that in the case of solidification for

traditional welding, subsequent to the sub-cooling, there

exists a thermal stabilization (columnar growth) and,

afterwards, if the conditions are present, there may be

equiaxial grain growth due to the constitutional sub-

cooling. However, in the case of welding via IEA, there is

a sub-cooling of the liquid, then a recalescence and,

finally, a continuous cooling at a high-cooling rate.

Consequently, it is important to evaluate the

temperature variation of the molten weld pool GL with

respect to the distance from the heat source, corresponding

to the fusion limit (solid–liquid) defined by

GL ¼dTL

dx: ð3Þ

Based on the graph measuring the temperature in the

fusion pool, it is possible to obtain the curve segment

corresponding to the cooling of the liquid welding metal

(Figure 11).

It can be seen that the transformation from liquid to

solid, during the solidification process, occurs in a

continuous manner from the fusion temperature of the

alloy of the filler material (6308C) to the solidification of

the latter (5818C).

Based on the polynomial approximation obtained in

the graph of Figure 11, it is possible to carry out the

calculation corresponding to the temperature variation in

Figure 9. Microstructure of the weld metal in a joint preheatedto 1508C, showing heterogeneous nucleation.

Figure 10. Comparison of cooling patterns between traditionaland IEA welding13.

Figure 11. Fraction of the cooling curve of the weld pool.

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the centre of the molten pool, in relation to the

advancement of the heat source, in the following manner:

GL ¼d

dxð41805:2 2 909:55xþ 5:016x2Þ

¼ 2909:55 þ 10:03x: ð4Þ

From Figure 11, the distance, 6 mm, travelled by the

heat source from a temperature of 811.9 to 5818C can be

obtained. Consequently, a temperature gradient of

849.378C/mm can be generated. In accordance with the

microstructural variation diagram developed by Kurz and

Fisher24, the type of microstructure in the centre of the

weld corresponds to a cellular dendritic type (Figure 9).

4. Conclusions

. The cooling rate at any condition of MIEA welding

exceeds the TTT curve of precipitate formation, b 0,

and, at a given time, these speeds tend to become

similar, since the mechanical resistance is practi-

cally the same due to the transition of precipitates of

the b 00 phase to b 0.. The incoherence of phase b 0 with the matrix reduces

the degree of hardening within the HAZ, and the

failure zone, after the traction tests, depends on the

preheating in MIEA welds.. The measurement of temperature in the molten pool

made it possible to determine the cooling charac-

teristics in the liquid metal and to calculate,

approximately, the type of microstructure expected

in the welding metal following solidification.

Acknowledgements

The authors would like to thank the National Council for Scienceand Technology for its support, the Universidad Michoacana deSan Nicolas de Hidalgo and, in particular, the TechnologicalInstitute of Morelia for all the facilities loaned for the use of itslaboratories.

Notes

1. Email: ricraf74@gmail.com2. Email: gbarrera@zeus.umich.mx3. Email: rgarcia@zeus.umich.mx

References

1. Ambriz RR, Barrera G, Garcıa R. S&I. 2006;11:10–17.2. Malin V. Weld J. 1995;74:305s–318s.3. Li Y, Murr LE, McClure JC. Mat Sci Eng A.

1999;271:213–223.4. Anderson NA. Instrumentation for process measurement and

control. p. 131–133.5. ASM. Heat treater’s guide: nonferrous alloys. p. 204.6. Huang C, Kou S. Weld J. 2004;83:111s–122s.7. Myhr OR, Grong O, Fjaer HG, Marioara CD. Acta Mater.

2004;52:4997–5008.8. Lu MJ, Kou S. Weld J. 1989;68:382-s–388-s.9. Hatch JE. Aluminum properties and physical metallurgy.

Metals Park, OH: American Society for Metals; 1993.p. 134–148.

10. Davies GJ, Garland JG. Int Metall Rev. 1975;20:83–106.11. Garcia R, Lopez VH. J Mater Sci. 2006;42:7856–7963.12. Garcia R, Lopez VH, Kennedy AR, Arias G. J Mater Sci.

2007;42:7794–7800.13. Garcia R, Lopez VH, Bedolla E, Manzano A. J Mater Sci

Lett. 2002;21:1965–1967.14. Garcia R, Lopez VH, Bedolla E, Manzano A. J Mater Sci.

2003;38:2771–2779.15. Lopez VH, Garcıa R, Bedolla E. Metall Mater Trans B.

2002;33:932–937.16. ASTM. Standard test methods of tension testing wrought and

cast aluminum- and magnesium-alloy products [Metric].West Conshohocken, PA: ASTM; 1994. p. 419–429.

17. Grong OO. Metallurgical modelling of welding. London:The Institute of Materials; 1997. p. 221–278.

18. Rosenthal D. Trans ASME. 1946;68:849–866.19. Sheppard T. Mater Sci Tech. 1988;4:636.20. Dutta I, Allen SM. J Mater Sci Lett. 1991;10:323–326.21. Kluken AO, Bjorneklett B. Weld J. 1997;76:39–44.22. Guiterrez LA, Neye G, Zschech E. Weld J. 1996;116s–121s.23. David SA, Vitek JM. Mat Rev. 1989;34:213–245.24. Kurz W, Fisher DJ. Fundamentals of solidification. Switzer-

land: Aedermannsdorf; 1989.

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