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Journal of the Brazilian Society ofMechanical Sciences and Engineering ISSN 1678-5878Volume 39Number 10 J Braz. Soc. Mech. Sci. Eng. (2017)39:4059-4068DOI 10.1007/s40430-017-0864-z

Formability of a wire arc depositedaluminium alloy

C. M. A. Silva, I. M. F. Bragança,A. Cabrita, L. Quintino &P. A. F. Martins

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TECHNICAL PAPER

Formability of a wire arc deposited aluminium alloy

C. M. A. Silva1 • I. M. F. Braganca2 • A. Cabrita1 • L. Quintino1 • P. A. F. Martins1

Received: 21 January 2017 / Accepted: 17 July 2017 / Published online: 26 July 2017

� The Brazilian Society of Mechanical Sciences and Engineering 2017

Abstract This paper is focused on the formability of the

aluminium alloy AA5083 deposited by wire arc additive

manufacturing (WAAM). The presentation draws from

metal deposition with a robotic welding system to

mechanical and formability characterization by means of

standard test specimens. Finite element analysis using

porous metal plasticity is utilized to model strain hard-

ening and the changes in porosity due to plastic defor-

mation. Results show that the deposited aluminium alloy

has excellent ductility and that its final stress response can

significantly improve as a result of strain hardening.

Voids resulting from metal deposition are closed by

negative values of stress-triaxiality resulting from com-

pression forming. The investigation is also a step towards

understanding the potential of including intermediate

forming operations in conventional wire arc additive

manufacturing (WAAM), consisting of metal deposition

and machining.

Keywords Formability � Aluminium alloy � Wire arc

deposited � Experimentation � Finite element method

1 Introduction

Recent years saw the development of innovative additive

manufacturing technologies to fabricate and repair metal

parts, faster and cheaper. Electron beam additive manu-

facturing [1], laser sintering additive manufacturing [2] and

arc-based additive manufacturing [3] are among the most

widespread additive manufacturing technologies that

combine the flexibility of metal deposition by welding with

the precision of CNC machining to fabricate prototypes

and production parts in low-batch sizes with less material

waste and shorter time-to-market than conventional man-

ufacturing technologies.

Arc-based additive manufacturing is the oldest of these

technologies and offers the advantage of using commonly

available equipment in metal working companies (e.g.

standard robotic MIG/MAG, TIG or plasma welding sys-

tems) instead of requiring the acquisition of a specific

additive manufacturing machine. Another benefit of arc-

based additive manufacturing is the high material deposi-

tion rate because it uses welding wire as feedstock instead

of using powder bed of fed-based systems as, for example,

laser sintering additive manufacturing.

The key topics in wire arc additive manufacturing

(WAAM) are the geometric sub structuring deposition

strategy and the selection and control of the arc-based

process parameters to ensure that metal is deposited, layer

by layer, accurately and free from defects until the part

Technical Editor: Marcio Bacci da Silva.

& P. A. F. Martins

pmartins@ist.utl.pt

C. M. A. Silva

carlos.alves.silva@ist.utl.pt

I. M. F. Braganca

ibraganca@dem.isel.pt

A. Cabrita

andre.cabrita@ist.utl.pt

L. Quintino

lquintino@ist.utl.pt

1 IDMEC, Instituto Superior Tecnico, Universidade de Lisboa,

Av. Rovisco Pais, 1049-001 Lisbon, Portugal

2 ISEL, Instituto Superior de Engenharia de Lisboa, Instituto

Politecnico de Lisboa, Rua Conselheiro Emıdio Navarro,

1959-007 Lisbon, Portugal

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DOI 10.1007/s40430-017-0864-z

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reaches near-net shape and is ready for the final finishing

by machining. In case of aluminium alloys, which will be

used in this investigation, major defects are caused by

porosities [4], and some authors propose the utilization of a

roller behind the welding arc to plastically deform and

close the voids by application of a vertical compression

force [5]. Inter-layer rolling, as this procedure is desig-

nated, also promotes microstructure grain refinement and

prevents distortion by significantly reducing the residual

stresses arising from metal deposition [6, 7]. The

improvement of the mechanical properties by inter-layer

rolling is also very important when the fully finished part is

machined from the deposited metal without any heat

treatment.

But what if additive manufacturing is extended to

include forming operations? The possibility of this hap-

pening is strong and should be considered at two different

levels:

• The basic level (Fig. 1a) is the case of metal parts

having cross-sections along different planes. This is

because fabrication by additive manufacturing will be

faster and cheaper (due to less removing of unwanted

material by machining) if forming operations (such as

bending) are included to distribute the deposited

material along the different required planes;

• The complex level (Fig. 1b) is the case of deposited

metals being used as preforms with optimized geometry

to ensure defect-free metal flow and complete die fill

with minor material losses during small batch forming

runs. In fact, the extension of additive manufacturing to

include forming operations such as forging offers the

advantage of reducing manufacturing time and tooling

costs while eliminating metal porosity, exceeding

wrought mechanical properties through strain harden-

ing and reducing final finishing by machining.

The last vision is not very much different from what is

commonly done in powder forging, in which compacted

and sintered metal powder preforms with a geometry close

to the final part are forged to produce a fully finished part

with residual porosity and large savings in material and

tooling costs [8].

Under these circumstances, the aim of this work is to

investigate the formability limits of the deposited metal due

to its importance for quantifying the resistance to initiation

and propagation of cracks during a forming operation. The

presentation starts by determining the stress–strain curve

along directions parallel and perpendicular to the metal

deposition path and proceeds towards the characterization

of the formability limits by fracture and to the validation of

these limits by means of a simple compression forming

process. The work is supported by experimentation and

finite element modelling.

2 Experimentation

2.1 Welding parameters

A Kuka 6-axis robotic system using a Fronius CMT VR

7000 cold metal transfer welding machine was utilized for

metal deposition. The machine works in controlled dip

transfer mode with a lower heat input than that of con-

ventional machines operating in dip transfer mode, and is

capable of delivering beads with excellent quality without

spatter. Computer interfaces are utilized to control the

movement of the welding torch and the welding

parameters.

The wire electrode was aluminium AA5083 with 1 mm

diameter, and the shielding gas was 99.9% of Argon.

Table 1 lists the major welding parameters.

The deposited metal sample consisted of a paral-

lelepiped with 50 mm 9 66 mm 9 180 mm (Fig. 2a) built

through the deposition of multi-bead overlapping layers

with a bead width w = 6 mm and a distance d = 4 mm

between the centres of adjacent beads (Fig. 2b). The ratio

d/dw = 0.67 was selected in accordance to the flat-top

overlapping model [9] to ensure an optimal overlap of the

weld beads.

The welding torch was programmed with a ‘zigzag’

movement (Fig. 2c), and its working angle was set to 45�.The travel angle was set to 08 to avoid differences between

forehand welding in one direction and backhand welding in

the other direction of the ‘zig-zag’ movement.

2.2 Mechanical characterization of the deposited

material

2.2.1 Stress–strain curve

The stress–strain curve of the deposited aluminium alloy

AA5083 was performed by means of standard compression

tests. The cylinder test specimens with 15 mm diameter

and 15 mm height were machined out from the sample

shown in Fig. 2a along the width ‘W’, longitudinal ‘L’ and

thickness ‘T’ directions, with the material in the ‘as-de-

posited’ condition (Fig. 3).

The tests were performed at room temperature on a

hydraulic testing machine (Instron SATEC 1200kN) with a

crosshead speed equal to 10 mm/min. The compression

platens were cleaned with ethanol before each experiment.

The cylinder test specimens were lubricated with Teflon

sheets on the top and bottom ends before compression to

reduce friction. Two different test specimens were utilized

for each ‘W’, ‘L’ and ‘T’ directions.

Figure 4a shows the average stress–strain curves

obtained for each direction. As seen, there are no

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significant differences in the stress response of the material

along the three different directions (‘W’, ‘L’ and ‘T’). This

means that from a macroscopic point of view, the

mechanical behaviour of the deposited aluminium alloy

AA5083 may be considered isotropic.

The average stress–strain curve resulting from the entire

set of experimental data is approximated by the following

Ludwik–Hollomon’s equation,

r ¼ 475:5 e0:22 ð1Þ

Fig. 1 Extending additive

manufacture to include forming

operations. a Basic level in

which a bending operation (left)

is included to diminish the

removing of unwanted

deposited material (right);

b complex level in which

deposited material is utilized to

fabricate preforms with

optimized geometry for small

batch forming runs

Table 1 Major welding parameters

Current (A) Voltage (V) Wire feed rate (m/min) Welding speed (mm/min) Stick-out length (mm) Gas flow rate (l/min)

100 16.5 8.5 600 12–15 17

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2.2.2 Friction

Friction at the contact interface between the compression

platens and the formability test specimens that will be used

in the investigation was estimated by means of ring com-

pression tests. The ring test specimens with a ratio of the

outer diameter, the inner diameter and the thickness equal

to 6:3:2 were machined out from the deposited aluminium

alloy AA5083 sample shown in Fig. 2a. Two different

lubrication conditions were evaluated: (1) dry friction and

(2) lubrication with molybdenum disulphide MoS2.

Friction was characterized by means of the law of

constant friction sf = mk, where sf is the friction shear

stress, m is the friction factor and k is the shear yield stress.

Figure 4b allows estimating m for both lubrication condi-

tions (m = 0.3 for dry conditions and m = 0.15 for

molybdenum disulphide MoS2) by comparing experimental

measurements and calibration curves relating the changes

of minimum internal diameter as a function of the reduc-

tion in height that were previously obtained by finite ele-

ment modelling.

2.2.3 Density

The porosity of the deposited aluminium alloy AA5083

sample was evaluated by machining out cylinder test

specimens along the width ‘W’, longitudinal ‘L’, and

thickness ‘T’ directions, and measuring its densities in a

laboratory precision balance equipped with a density

determination kit.

The experimental technique utilized in the determina-

tion of density was based on the Archimedes’ principle (or

buoyancy) method, commonly utilized in powder forging,

which states that a body immersed in a fluid loses weight

by an amount equal to the weight of the fluid it displaces. A

sample of the electrode wire was utilized for determining

the density of the fully dense aluminium alloy AA5083 for

reference purposes. The results are provided in Table 2.

Distilled water at 22 �C was utilized in the determina-

tion of densities.

By defining the initial uniform relative density R0 as

follows,

Fig. 2 Deposited aluminium

alloy AA5083. a Photograph of

the sample; b schematic

diagram of the overlapping

between layers of the weld

beads; c schematic ‘zigzag’

movement of the welding torch

Fig. 3 Compression test

specimens. a Schematic

representation of the deposited

metal sample with the width W,

longitudinal L and thickness

T directions; b photograph of a

cylinder test specimen

machined out from the

deposited aluminium alloy

AA5083 sample

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R0 ¼ qsamples

�qwire ð2Þ

where qsamples is the average density of the test samples and

qwire is the density of the fully dense wire, one obtains

R0 = 0.997. This result allows concluding that the average

porosity of the deposited aluminium alloy AA5083 is

approximately equal to 0.3%.

The above procedure allows quantifying the percentage

of average porosity in volume in an easier and more

effective way than that based in two-dimensional images

taken from selected cross-sections of the test samples.

2.3 Formability of the deposited material

Conventional bulk formability tests performed with cylin-

drical and tapered specimens compressed between flat

parallel platens were utilized to determine the onset of

failure by fracture of the deposited aluminium alloy

AA5083 sample. The formability tests were performed on

the same hydraulic testing machine that had been previ-

ously utilized for the mechanical characterization of the

material. The experiments were carried out at room tem-

perature with a crosshead velocity of 10 mm/min.

The formability test specimens were machined out from

the deposited sample shown in Fig. 2a, according to the

geometries provided in Table 3. The lubrication conditions

utilized in the tests are also provided in Table 3.

Fig. 4 Mechanical and friction characterization of the deposited

aluminium alloy AA5083. a Stress–strain curves for test specimens

taken from the width W, longitudinal L and thickness T directions;

b experimental measurements and ring test calibration curves relating

the changes of the minimum internal diameter with the reduction in

height for several friction factors

Table 2 Density of the electrode wire made from fully dense alu-

minium alloy AA5083 and average density of the test samples cut out

from the deposited aluminium alloy AA5083

Wire Test samples(Average values)

Mass out of wate r (g) 2.062 11.820Mass in water (g) 1.279 7.319

Volume (cm3) 0.785 4.513

Density ρ (g/cm3) 2.626 2.619

Table 3 Geometry and lubrication conditions utilized in the forma-

bility tests

Cylinder specimen Tapered specimen

H

D

w0

h0

D

d

tH

H (mm) 25 25

D (mm) 25 30

endD (mm) – 25

t (mm) – 5

Lubrication Dry MoS2 Dry

Label Cd Cl Td

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The procedure utilized to determine the experimental

strains at fracture and the corresponding fracture locus was

similar to that described by Silva et al. [10]. The procedure

makes use of rectangular grids with 3 mm initial side

length that are engraved at the mid-height of the specimens

(refer to the black squares in Table 3) and measured for

different vertical displacements of the compression platens

in a Mitutoyo TM-111 toolmaker’s microscope equipped

with an XY measurement micrometre table.

The experimental axial ez and circumferential eh strains,corresponding to each vertical displacement of the com-

pression platens in which grids are measured, are obtained

as follows,

ez ¼ lnh

h0eh ¼ ln

w

w0

ð3Þ

where h0 and h are the initial and actual heights, and w0 and

w are the initial and actual widths of the rectangular grid.

The radial strain is determined by incompressibility

er = -(ez ? eh).

3 Finite element modelling

The experimental tests were simulated with the in-house

finite element computer program I-form that is being

developed and validated against experimental results since

the end of the 1980s [11]. I-form allows modelling the

plastic deformation of porous metals by means of the fol-

lowing functional,

P ¼Z

V

rij _eij dV �Z

ST

Ti ui dS ¼ 0 ð4Þ

where rij and _eij are the stress and strain rate tensors, and Tiand ui are the surface tractions and velocities on the surface

ST.

The constitutive equations of the yield criterion utilized

for modelling the plastic deformation of porous metals

proposed by Doraivelu et al. [12] are written as follows,

rij ¼�r_�e

2

A_e0ij þ

dij3 ð3� AÞ _ev

� �ð5Þ

where the symbols �r and _�e denote the effective stress and

effective strain rate, r0ij and _e0ij represent the deviatoric

stress and deviatoric strain rate, and _ev is the volumetric

strain rate given by _ev ¼ _R�R, where R is the relative

density. The initial uniform relative density R0 = 0.997 of

the deposited aluminium alloy AA5083 was experimen-

tally determined by means of Eq. (2).

The effective stress �r in Eq. (5) is a function of the first

invariant of the stress tensor J1 and of the second invariant

of the deviatoric stress tensor J02,

�r2 ¼ AJ02 þ BJ1

A ¼ 2þ R2 B ¼ 1� A

3C ¼ 2R2 � 1 ð6Þ

In what regards modelling the onset of failure by frac-

ture, it was decided to utilize an uncoupled ductile damage

approach that made use of the normalized version of the

Cockcroft and Latham [13] criterion,

Dcrit ¼Z�ef

0

r1�rd�e ð7Þ

where r1 is the major principal stress and �ef is the effectivestrain at fracture. The choice of Cockcroft and Latham’s cri-

terion instead of another ductile damage criteriawas due to the

fact that typical crack opening mode in upset compression

processes (bulk forming) is by ‘out-of-plane shearing’, also

known as ‘mode III’ of fracture mechanics [14].

The finite element models of the cylindrical and tapered

formability test specimens took advantage of the rotational

symmetry conditions of the experimental setup. The

material was discretized by means of linear quadrilateral

elements with an initial uniform relative density

R0 = 0.997, to account for the porosity of the deposited

aluminium alloy AA5083 sample. The compression platens

were modelled as rigid objects and were discretized by

means of linear contact elements with friction. Figure 5

shows the initial and final deformed mesh of the tapered

formability test specimen.

The numerical simulations were carried out through a

succession of displacement increments each one modelling

approximately 0.1% of the initial height of the specimen

and the overall computing time for a typical analysis

Fig. 5 Initial and final deformed mesh of the finite element model

utilized for the numerical simulation of a tapered test specimen

compressed between flat parallel platens

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containing 6000 elements and 6100 nodal points was below

1 min on a standard laptop computer equipped with an

Intel i7 CPU (2.7 GHz) processor.

4 Results and discussion

4.1 Formability tests

Figure 6a, b show the finite element predicted geometries

of the cylindrical (‘Cd’) and tapered (Td’) formability test

specimens (Table 3) at the onset of failure by fracture and

the corresponding distributions of effective strain �e, relativedensity R and accumulated ductile damage D.

As seen, apart from a small localized region of the

specimens containing their mid-height free surfaces where

the relative density R % R0 % 0.997 is similar to that of

the original deposited metal, the remaining regions are

fully dense R % 1. This is due to high negative stress-

triaxiality rm=�r resulting from the compression loading

applied by the platens and allows concluding that inter-

layer rolling during metal deposition is not needed in case

Fig. 6 Formability of the deposited aluminium alloy AA5083.

a Finite element computed distributions of effective strain for the

formability test specimens ‘Cd’ (left) and ‘Td’ (right) of Table 3;

b finite element computed distributions of relative density (left part of

the distributions) and ductile damage (right part of the distributions)

for the formability test specimens ‘Cd’ (left) and ‘Td’ (right) of

Table 3; c experimental and finite element predicted evolutions of the

force with displacement for test specimens ‘Cd’ (left) and ‘Td’ (right)

of Table 3

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forging operations are included in the additive manufac-

turing route. In fact, not only voids are closed as material

strain hardening (refer to the distributions of effective

strain in Fig. 5a) but it also significantly improves the

overall stress response of the final part.

The finite element results and inserts photographs in

Fig. 6a, b also allow concluding that failure by cracking is

triggered at the mid-height free surface of the specimens, in

close accordance with the distribution of the accumulated

damage given by the normalized version of the Cockcroft–

Latham ductile damage criterion (7). In fact, contrary to

what one might think, the out-of-plane shear cracks that are

observed in both specimens are caused by the accumulation

of ductile damage beyond a critical value D[Dcrit and not

by the existence of residual porosity.

The vertical displacement of the compression platens

corresponding to the instant of time where cracks are

triggered coincides with the end of the experimental evo-

lutions of the force with displacement in Fig. 6c. The

overall agreement between the experimental and finite

element predicted evolutions is very good.

The experimental and finite element predicted strain

loading paths for the three different formability testing

conditions of Table 3 are shown in Fig. 7. As seen, the

limiting fracture strain pairs on the free surface of the spec-

imens where cracks are triggered fall on a fracture locus

consisting of a straight linewith slope ‘-1/2’. This result is in

excellent agreement with what is commonly found in spec-

imens made from wrought (fully dense) ductile metals.

In fact, investigations by Kudo and Aoi [15] and Kuhn

et al. [8] performed in the late 1960s and early 1970s

showed that the limiting fracture strain pairs on the outside

surfaces of upset test specimens made from wrought duc-

tile metals fall on a straight line with slope ‘-1/2’ in the

principal strain space (that is, fall on a line parallel to

uniaxial compression loading under frictionless

conditions). Martins et al. [14] recently proved that the

fracture loci of wrought ductile metals with slope ‘-1/2’

are associated to failure under crack opening mode III.

As a result of this, it can be concluded that the slope and

the crack opening mode (by out-of-plane shear stresses) of

the deposited aluminium alloy AA5083 are identical to

those commonly observed in wrought ductile metals.

Moreover, the y-intercept where the experimental straight

line with slope ‘-1/2’ of the fracture locus crosses the eh-axis, indicates good ductility of the deposited aluminium

alloy AA5083.

4.2 Compression forming

Heading is a compression forging operation for enlarging

and reshaping some of the cross-sectional area of a rod. In

its simplest form, the deformation is accomplished by

holding the rod between grooved dies and applying axial

compression to its end by means of a flat compression

platen.

The advantage of modifying the conventional additive

manufacturing route consisting of metal deposition and

machining to include an intermediate heading stage is

related to the benefits in material and time deposition

savings and in scrap removal by final machining, due to the

possibility of spreading the end of the deposited metal

stock by plastic deformation. This is interesting for small

production runs of flanged components, among other parts.

As a result of what was mentioned above, the aim of this

last section of the paper is to evaluate the overall forma-

bility of the deposited aluminium alloy AA5083 during a

typical cold heading operation and to analyse the distri-

bution of porosity in the final part.

Figure 8a, b shows the initial and finite element pre-

dicted geometries of the cold headed part and the computed

distribution of effective strain, relative density and ductile

Fig. 7 Fracture locus of the

deposited aluminium alloy

AA5083 in the principle strain

space determined from the

experimental fracture strains of

the formability test specimens

listed in Table 3. The solid

markers correspond to

experimental strain pairs

measured for several

intermediate levels of

deformation and the lines are

the finite element computed

strain loading paths

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damage. As seen, plastic deformation and material strain

hardening are localized in the upper end of the rod. The

amount of accumulated ductile damage after which cracks

are triggered (Dcrit % 0.2) is compatible with that of the

formability test specimens, and the corresponding fracture

strains and strain loading paths are also in good agreement

with the previously determined fracture locus (Fig. 7).

The computed distribution of relative density is also

compatible with the experimental observations of porosity

in different regions of the cold headed part (please refer to

the inset photographs). This result allows concluding that

the extension of additive manufacturing to include heading

avoids the utilization of inter-layer rolling because voids

are closed by the high negative values of stress-triaxiality

rm=�r. The final strength is also significantly increased by

material strain hardening.

5 Conclusions

The mechanical and formability characteristics of an alu-

minium alloy AA5083 sample deposited by wire arc

additive manufacturing were investigated by means of

conventional tests performed on cylindrical and tapered

test specimens. Results show that the deposited aluminium

alloy AA5083 is isotropic and that its overall ductility is

appropriate to extend conventional additive manufacturing

routes to include intermediate forming operations.

The results from experimental and finite element anal-

ysis of formability revealed that the limiting strain pairs of

cylindrical and tapered test specimens made from the

deposited aluminium alloy AA5083 fall on a fracture locus

consisting of a straight line with slope ‘-1/2’ in the prin-

cipal strain space. This result is similar to what is

Fig. 8 Cold heading of the deposited aluminium alloy AA5083. a Finite element computed distributions of effective strain, relative density and

ductile damage at the instant of deformation where cracks are triggered; b computed strain loading path in the principal strain space

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commonly found in specimens made from wrought (fully

dense) ductile metals, and is typical of crack opening by

out-of-plane shear (mode III of fracture mechanics).

Experimental and finite element modelling of the cold

heading of a rod made from the deposited aluminium alloy

AA5083 not only confirmed the formability limits deter-

mined with cylindrical and tapered test specimens as they

also confirmed that intermediate forming operations with

compressive-dominant stress states successfully promote

the closure of voids and improve material strength through

strain hardening. This last result allows circumventing the

need of inter-layer rolling to eliminate porosity.

Acknowledgements The authors would like to acknowledge the

support provided by Fundacao para a Ciencia e a Tecnologia of

Portugal and IDMEC under LAETA—UID/EMS/50022/2013 and by

the European Commission under the H2020 Programme within the

project Lasimm ‘‘Large Additive Subtractive Integrated Modular

Machine’’—Grant Agreement 723600.

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