AIP Conference Proceedings 2113 , 030001 (2019); https://doi.org/10.1063/1.5112529 2113 , 030001
© 2019 Author(s).
Backward rod extrusion of bimetallic
aluminum-copper alloys at room
temperature
Cite as: AIP Conference Proceedings 2113 , 030001 (2019); https://doi.org/10.1063/1.5112529
Published Online: 02 July 2019
V. Sanabria, S. Gall, F. Gensch, R. Nitschke, and S. Mueller
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Backward Rod Extrusion of Bimetallic Aluminu m -Copper
Alloys at Room Temperature
V. Sanabria 1, a) , S. Gall 1, b) , F. Gensch 2, c) , R. Nitsch ke 2, d) , S. Muelle r 2 , e)
1 INGWERK Gmb H. Gustav -Meyer-Allee 2 5, Gebäud e 17a, Treppe 5, 133 55 Berlin.

2 Extrusion Resea rch and Development Center ERDC, TU Berlin. Forschun gszentrum Stran gpressen, Sek r. TIB 4/1 -
2. Gu stav-Meyer-Allee 2 5, 13355 Berlin

a) Corresponding aut hor: vidal. sanabria@in gwerk.com
b) sven.gall@ing werk.co m
c) felix.gensch@stra ngpressen.b erlin
d) rene.nitschke @strangpressen.b erlin
e) soeren. mu eller @strangpress en.berlin

Abstract. The backward rod extrusion of bimetallic aluminum-cop per alloys at roo m temperature was in vestig ated. The
aluminum allo y EN AW-10 80A and the copper alloy Cu -ETP were selected to p repare th e core and sleeve of the billet
respectively . The co pper cross se ction was equ ivalent to 30% o f the b illet. Moreover, the billet was extruded applying a
conic die angle of 90° and an extrusion rati o of 1 4:1 . Experim ental results dem onstrated that the com bination of gro unding
marks on the die surfac e and th e application of graphit e foil reduced d rastically the friction betwe en copper and the conic
die. Thus, a uniform material flow o f aluminum and copper through the bearing ch annel was observed during the steady
state of th e extrusion process. How ever multiple fractures of the copper sleeve occu rred at the end of the process. The
extrusion pro cess was nu merica ll y si mulated applying the FEM -based software Defor m 2 D in order to estimate the state
variables and ma terial flo w . The d ie and punch temperature evoluti on, as well as the d ie extrusion force were recorded
during the whole process to facilitate the validation of the nume rical analysis.
INTRODUCTION
The excellent electr ical and t hermal co nductivity of co oper alloys make them ideal materials for conducto rs.
Ho w ever, the high density, lo w availabilit y and instable prices have m otiva ted their substitution in industrial
applicatio ns. Due to the r ich availabilit y, lo w p rice and density, as well as goo d electric conductivity of al uminum
allo y s, the y have bee n used a s a substitu te material for co pper for more than fifty years. Nevert heless, a hard and
brittle aluminum oxide appears o n the surface increa sing thus its electrical contact resis tance. Copp er -clad alu m inu m
compound is an efficie nt altern ative for this c hallenge and gives a goo d co m pro m ise bet ween conductiv ity and weight.
Thus, the co m pou nd weight ca n be r educed in 50 % for an equivale nt co nductivity compari ng to copper and
additionall y a b etter contact can be achieved. Cop per-clad aluminum pro files ar e usually manufactured by means of
hydrostatic extrusion (Fi g. 1 (a)) .

1

Aluminum al loy

2

Copper al loy

3

Fluid

4

Die

5

Container

6

Punch

α

Half die a ngle (° )

(a) (b) ( c)

FIGURE 1. Schematic represe ntation of (a) hy drostatic, (b) direct and (c) indirect extrusion.

Pr oceedings of the 22nd International ESAFORM Confer ence on Material F orming
AIP Conf. Proc. 2113, 030001-1–030001-6; https://doi.org/10.1063/1.5112529
Published b y AIP Publishing. 978-0-7354-1847-9/$30.00
030001-1

During this process a fluid transmits the driving force from punch to w orkpiece, which is f orced to flo w plasticall y
through the die. Mo reover, th e fluid acts also as lubrica nt re ducing significa ntly t he friction bet ween slee ve (cop per)
and d ie and therefore a m ore uniform m aterial flo w i s po ssible. On the other hand, the complicated b illet preparatio n
process, sealing issue as well as th e long cycle times m ake s the hy dro static extrusion a com plex and expensive proce ss.
In the for ward and bac kward extrusion the billet is i n co ntact with t he too ls a nd i mportant frictional and ther m al
effects are generated . Due to the hi gh friction between th e sleeve and the container/di e in the forward extrusion
(Fig. 1(b )) and sleeve and die in the backward extr usion (Fig. 1 (c)), the sleeve flo ws slo w er than the core. T his
phenomena generates non -unifor m m aterial f low and undesi red frac tures of the sleeve or core m aterial. Experi mental
investigatio ns have de m onstra ted that cop per clad aluminum profiles can be also manufactured b y means of forward
and back ward extrusion under deter m ined billet configuratio n and pro cess parameters. In g ene ral, author s agree that
a conic die, low extr usion ratio and a w ell lubr ication betwee n d ie and slee ve faci litate a so und material flow a nd
minimize the fractures [1 ,2,3]. A s m all conic die an gle as w e ll as a l ow extrusion ratio reduce th e normal contact force
bet w ee n the sleeve an d die. T his geo m etrica l condition coupled with an appropriate lu brication reduce s the friction on
the die sur face and therefore generate s a m ore unifor m material flo w [4] . Kang and K w o n success fu ll y e xtruded cop per
clad aluminum pro files with a maximal e xtrusion ratio o f 7:1 [ 1 ,2]. They e xtruded short pro files witho ut fracture using
a conic die angle of 60° w hich was lubricated w ith carbo n o il. Additionally, th e core and sleeve w ere heated to 300 °C
and 350 °C respectively be fore extruding.
A more i ntimate i nitial i nterface bonding bet ween co re and sleeve also i m prove the material flow and the final
mechanical and t hermal interface prop erties. Diff ere nt billet prep aration meth ods suc h as p ressurization j oining b y
cold co m pressio n [1], low pressure casting [ 3] and exp losive clad ding [5 ] of the aluminum ro d in the cop per pipe are
applied . However, these co m plex p reparatio n methods applied in conventional for w ard or back w ard e xtrusion make
the processes m ore e xpensive and red u ce s the ir competitive ness in relatio n to hydrostatic extrusion.
The initial billet temperature plays also an i mportant role in the appro priate material flow duri ng co pper clad
aluminum e xtrusion. Experimental and numerical i nvestigations s howed a more unifo r m material flow at lo wer
temperatures [ 2,3 ]. At h igher te mperatures the flow stres s d iff ere nce betwee n t he alumi num a nd cop per increases.
Kwon and co workers sug gested that t he flow stress o f co pper was 4 to 7 ti mes hi gher than t hat of p ure aluminum at
35 0 °C, b ut 7 to 13 ti m es higher at 4 70 °C [ 2]. T herefore, a longer al uminum nose cro p is usua lly o bserved when a
higher initial billet te mperature is app lied .
In this work industrial scaled back w ard extrusion of co pper clad alum in um p rofiles was investigated . For that
purpose, a si m ple billet prep aration w as carried out. A dditionall y, a relative high extrusion r atio (1 4 :1) and a moderate
conic die ang le (90 °) w a s ap plied. Moreover, th e billet an d th e tools w ere not heated an d th e deformation process w a s
performed initiall y at roo m te m perat ure.
EXPERIM ENTAL PROCEDURE
A b ackward rod extrusion o f bi m etallic al uminum -co pper alloys was car ried out at the ERDC of t he T U B erlin.
Since the 8MN p ress is equipped w ith load cells, the extrusion for ce w as reco rded d uring the whole pr ocess. The
aluminum allo y EN A W-10 80A and the cop per alloy Cu -ETP w ere selected to prepar e the co re and sle eve of t he billet
respectively ( Fig. 2 (a)) . The copper alloy was so ft an nealed at 600 °C for 3 hours in ord er to r educe the flow stress.
The core was intentionally shor ter than the sleeve to reduce the nose crop [ 6]. Moreover, the c ore and sleeve surfaces
were cleaned with ethanol before assem bl y . A clearance f it between them was applied to facilitate th e billet
prepar ation. Furthermore, the external diameter o f the core and sleeve were 77.4 mm and 93 mm, respecti vely. Thus,
the co pper cross section was eq uivalent to 30% o f th e billet section .

( a)

( b)

FIGURE 2. Extrusion asse mbly. ( a) Sch em atic representation of block assem bl y and (b) extrusion die.

030001-2

After core and slee ve were asse m bled, t he b illet was co m pl etel y co vered with the graphi te paper Gr afoil GTB
(250 µm th ick ness) wh ich i s usually used to reduce the fric tion bet w een workpiece and to ols during hot co mpression
test [7 ]. Subseq uently, the bil let was i ntroduced in to a con tainer ( Ø=95 mm) at roo m te m perat ure and indirec tly
extruded against a conic die (2α =90°) w it h a ra m speed of 5 mm/s . The initial tem perat ure of billet, container, die and
punch was 25 °C appro xim atel y. I t is important to notice t hat the conic die f ace w as not p olished bu t gro und with Si C-
paper grid 8 0 ( Fig. 2(b) ). In addition, the die -face was cove red with a graphite -based lubricant which was burned in
the surface at 400 °C. T hermocouples w ere p laced i n the die bearing c hannel and in the pu nch head in or der to register
the temperature evol ution (Fig . 2 (a)).
RESULTS AND DISCUSSI ON
The extrusio n of the Copper clad aluminu m profile at roo m te mperature c ould be car ried out appl y ing a co nic die
and a n extrusio n ratio o f 1 4:1. Figure 3 descr ibes t he m o st releva nt geo metric c haracteristics o f the pro file. The
extruded ro d had a full alu minum head (zo ne A) with a lengt h of 110 mm, follo wed b y 14 90 mm of unifor m aluminum -
copp er b im etallic co m pound (zone B ). In the last part of the ro d (zo ne C) multiple fractures o f t he cop per sleeve
occurred , and actually t his sec tion wa s longer (8 90 mm) tha n expec ted. T he nose crop (Fi g. 3, zone A) co uld not be
avoided even usin g a longer sleeve, mainly d ue to the greater flow resista nce of t he co pper. T his effect was also
observed in si milar investigatio ns [ 3,5,6 ] . Additionally, the sleeve fract ure has bee n also repo rted and it represents
one o f the big c hallenges in the co pper-clad aluminum e xtrusion. In Ref (2) it was suggested that the copp er fracture
occurs due to the large dif ference of flow stress bet w ee n core and sleeve materials.
In order to investigate the ther m o -mechanical respo nse of EN AW-108 0A and soft annealed Cu -ETP hot
compression test s were carried out. T hus , s m all speci m ens (diameter=10 mm, hei ght= 1 5 mm) were tested in the
Gleeble System 3800 at different te m perat ures (2 0 - 300 °C) and strain rates ( 0.01 - 10 1/s). Results of the test s are
depicted in Fig. 4. Fi gures 4 (a) and 4 (b) show strain- stress curve s of Cu-ET P and EN AW – 10 80A respective ly. As
expected, the flo w stress i ncreases with the strai n a nd strain rate b ut it is red uced at higher temperatures. Ho wever,
m ore rele vant tha n the magnitude of the flow stress is the relative value between flow stress of alumin um and copper .
Ho t compression tests de monstrated that when th e allo y EN AW-1080 A is h eated f ro m 20 °C to 300 °C its f low stress
is reduced m ore tha n 50 %, whereas the Cu -ET P alloy decre ased 30 % under the same conditions (F ig. 4(c)) . In other
words, the f lo w stress ratio σ Cu /σ Al i ncreases from 3.3 to 5 .1, co nsidering a strain rate of 10 1/s and strain 1 . It confirms
the observations made in [2] .
The extrusion proce ss was numericall y anal yzed app lying the FEM -based so ftware Defor m 2D in order to
investigate more abo ut the material flo w a nd state variables. Constitutive m odel s σ= f (ε, 𝜀 󰇗 ,T ) of the aluminum allo y
EN A W -1080A and th e copper alloy Cu-ET P w er e set based on th e experimental results of the hot co m pres sion tests.
Additionall y , the flow stress c urves were e xtrapolated from a strain 1 to 5 using expone ntial equations fitted to the
experimental results. T he T resca friction model was selected to simulate t he friction. Thus, a friction factor m =1 w a s
set b etween cor e and sleeve, a nd m =0 .17 between w or kpieces (co re and sleeve) and too ls. Friction factors of m =0.2
and m =0.2 5 have b een reported in extrusion trials where lubrication w ith carbo n oil has been ap plied [ 1 ,3]. T he heat
transfer co eff icie nt of 30 kW/ m 2 was set b etween all o bjects [ 1]. Additionally, the die, container and punch w ere
simulated as ri gid objec ts (no elastic deformation), but i nternal heat transfer was allowed.
Figure 5 (a) depicts the experimental and si mulated extrusion force (die force) during t he whole stroke ( 175 mm ).
Moreo ver, the si mulated material flow o f core and sleeve is shown in Fig. 5 (b). T hree main points of the c urve are
emphasized. T hus , point 1 indicate s the instant when the b earing channel is filled with alu minum to form the nose
crop (Fig. 3). Moreo ver, the point 2 s hows t he maximal extrusio n force, which is generated wh en the copp er sleeve
begins to flo w though the bearing cha nnel.

FIGURE 3. Description of the extrudate.

030001-3

( a)

( b)

( c)

FIGURE 4. Therm o- mechanical behavior of copper and alum inum alloys. (a) Strain -stress curves of homogenized Cu-ETP ,
(b) strain-stress curves of EN AW-1080 A and (c) variation of the flow stress ratio at high er tem peratures .

Additionall y , the p oint 3 illustrates the beginnin g o f multiple sleeve fractures also depicted as zone C (Fig. 3) .
Finally , the si mulated die force and the sequence o f the m at erial flo w are depicted in Fig. 5 (a) and 5 (b), resp ectively.
The numerical ana lysis co uld estimate the magnitude of the die force in the p oint 1 and in the stead y state (Zone B)
with an acceptable di fference of 5 %. Moreover , fracture co uld not b e pred icted, since fracture b ehavior was not
simulated. Ho w e ver the d ie force in the p oint 2 w as ar ound 25 % underesti mated. T he reason for this discrepanc y
could be related to an inappropriate simulation of the friction bet w ee n copp er and die (step 397). During the simulation
a constant friction factor m =0. 17 w as set b etween slee ve and die. It m eans t hat the normal force bet w een bo th metals
and its influence o n the fric tion b ehavior w as not co nsidered. In the real ca se, sleeve and di e surface could shortly
stick under ver y high nor mal pressure (s uch as t he p oint 2) par ticularly using the grounded d ie face ( e ven using
graphite paper as lu brication ) th en subsequently start to s lip. T his transient friction m ec hanism increases significant l y
the local shear stress and ther efore the to tal die force . As th e temperature increases in t he p rim ar y defor mation zone
and the co pper sleeve beco mes thinner the req uired extrusion force as well as t he normal force d rop. Under this contact
phenomena the f rictio n b ehavior chang es and it ca n be only par tially rep roduced applying a constant friction f actor m .
Figure 6(a ) depits the experimental and simulated geometry of the nose crop. It c an be observ ed that the le ngth of
the alumi num head was pro perly predicted with onl y 5 % of deviation. Nevert heless, the l ength of t he nec king zo ne
was ab out 30 % o verestimated. The pr ediction of the aluminum nose geometr y is stro ngly related to the friction
m odell ing b etween worpieces and die during the first 37 mm o f stroke (Fig. 5(b )). In this part o f the si mulation, the
conic die is filled by the core and sleeve materials, wh ich ar e force d to flo w plastical ly through the bearin g channel.

( a)

( b)

FIGURE 5 . Experime ntal and num erical results. (a) Co m parison of experimental and simulated die force and (b) simulated
material f lo w.

030001-4

(a)

( b)

FIGURE 6 . Experime ntal and num erical co m parison o f the initial m aterial flow. (a) Nose crop and (b) initiation of the
bimetallic ex trude (stroke 40 mm , step 431) . Extrusi on direction from left to right.
In ge neral, a high friction facto r b etw een die and sleeve delays t he cop per fl o w and gener ates a lon ger alu minu m
head. Since the head length was corr ectly p redicted, a good friction simulatio n u ntil the stro ke 35 mm ( step 3 62) can
be assumed (Fig. 5 (b)). On the other hand, t he inaccurate pred iction of the necki ng len gth sugges ts that t he same
friction relatio nship used until the step 36 2 are inapropiate for the steps 36 3 -397. Friction is influenced b y te mperature,
pressure, speed, etc and th erefore m ore accura te results can be only obtained applying a more sensitive f riction model
m= f (T ,P, v) [7,8]. In add ition, the lon ger necki ng legth a nd lo w er extr usion force (point 2 in Fig. 5(a)) obtained w ith
the numerical simulation ca n b e also explained d ue to an in sufficient metal bonding (lo w friction) b etween co re and
sleeve. A higher metal bo nding bet ween t he Cu -ETP and EN AW -108 0A alloys cou ld take p lace at t he bearin g channel
generating a m ore severe defor m ation and therefore a higher extru sion force at the beginning o f the extrusion p rocess .
Moreo ver, this stronger m eta l bonding could reduce the co re speed and thus the le ngth of the aluminium head. Fi g. 6 (b)
compares the experi mental and si mulated geometries at t he front end of the bimetallic ro d.
Initially, a ther m ocouple was placed inside the bear ing c hannel to meas ure the pr ofile t emperature ( Fig. 2 (a)) .
Ho w e ver, this thermoco uple was pushed backwards into the bore and covered by remaining graphite paper during the
extrusion . T hus, t he measured tempertaur e cor responds to the bearing land and not t he p rofile surface . Figure 7 (a)
compares the experimental an d simulated te mperature s of the die land and stem surface. Th e simulated die te mperature
was taken at a location 0.3 mm behind the bearing land. Experimental and simulated te mperatures are very si milar u p
to point 3 (stroke 130 mm ), where fracture too k place . Once the die force , m ater ial flo w and temperature evol ution
were validated until stroke 13 0 mm , simulat ion results ca n be used to analyse the o ccurrence of the fracture.

( b)

(a)

(c)

FIGURE 7 . Experime nt al and sim ulated results. (a) Experime ntal and estim ated temperature, (b ) sim u lated temperature
distribution in primary extrusion zone and (c) simulated effective strain in th e bea ring channel across the copper alloy.

030001-5

The exact origin o f the sleev e fracture in the zone C (Fig. 3) is still o bject of investigatio n, ho w e ver it ca n be
related to its te mperature ev ol ution, strain an d friction inside the bearing channel. A ccording to t he simulatio n r esults,
the cop per sleeve placed insi de the bearin g cha nnel is t he hottest area in t he p rimary d efor m atio n zone . I t reached
around 18 0 °C at the stro ke 40 mm (step 397 ) an d th en 250 °C at the fracture instant (Fig. 7 (b)). A higher temperature
of the slee ve can be related to the fract ure i nitiation, ho wever a hi gher fracture elo ngation i s allo wed at higher
temperatures [9 ]. For that reason, a te mperature increase dur ing the extru sion p rocess may not cause this fa ilure. On
the other ha nd, a temperat ure increase ge nerates a hi gh er flo w stress r atio σ Cu /σ Al as sho wn i n Fig. 4( c), what ma y
initiate fracture [2]. Mo reover, d uctile fracture inve stigation s of Cu -ET P revealed the significant influence o f the strai n
on the formation and nu cleatio n of voids [ 10 ]. Thus, the fracture w il l be initiated at a critical plastic strain.
Theor etically , during the backward extrusio n the w or kpiece experiences an average strain in th e primar y deformation
zone, which r emains constant after reaching the stead y state stage [ 11 ]. Numerical si mulation m ade i n the present
investigatio n revealed that strain cha nged due to the setti ng up or compression of t he billet in the die at the beginning
of the extr usion, as well a s the thickness variatio n of the cop per sleeve . T hus, the sleeve t hickness inside the bearin g
channel was for example 5.7 mm and 2.1 mm at the stroke 40 mm and 130 mm, resp ectively. Figure 7© sho ws the
effective strain ac ross t he co pper sleeve and its thic kn e ss. T he nu merical res ults p redict ed a maximal e ffective strain
of 2.7 at the strokes 40 mm a nd 3.8 at the fracture instance ( stroke 130 mm ) . Another poss ible cause for the fracture
occurrence is the strai n may increase due to red uction of lu brication bet ween sleeve and die . In the si mulation a
constant friction factor w as applied , how ever the lubrica tion capacit y can be pro gressively reduced d uring the
extrusion stro ke.
SUMMARY
An indu strial scale b ackward extrusion o f co pper clad aluminum p rofile was in vestigated at roo m temperature.
Aluminum allo y E N AW-1080 A and the cop per allo y Cu -ETP w ere app lied in the co re and sleeve o f t he bille t
respectively. Almost 15 00 mm o f bi m etallic pro file was su ccessfully extruded , mainly due to the friction reduction
bet w ee n the copp er sleeve and the co nic die face. This w as possible since t he billet was co mpletel y co vered w ith the
graphite p aper Grafoil GT B (25 0 µm thicknes s). Additio nally, the co nic die face was with a hig h surface roug hness
by grinding with SiC -paper grid 80 an d s ubsequentl y co vered w ith a graphite -based lubrica nt, w hich was b urned onto
the surface at 400 °C.
Fracture of the copp er sleeve was observed at 75% o f the ste m stroke (130 mm ) . The exac t origin of this failure is
still under investigation. Moreo ver, the increase of the flo w stress ratio σ Cu /σ Al at higher t emperatures, as well as the
progressive increase o f the plastic strai n seem to b e the main cause for the fracture occurrence.
The investigation sho ws that a co ld aluminum co re and ho t copp er sleeve would r educe the flo w stress ratio
bet w ee n the two app lied allo ys and could co ntribute to a more uniform material flo w .
ACKNOWLEDGEMENTS
The authors are grateful for the fin a ncial suppo rt of the German Research Foundation (D FG) u nder the co ntract
no. MU296 3/14- 1.
REFERENCES
1. C. Kang and H. K won, Internatio nal Journal of M echanical Sciences 44 , 247 -267 (2002).
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3. H. Kwon, T . Jung , S. Lim and M. Ki m, Materials Science Fo rum 449 - 452 , 317-320 (2004).
4. B. Avitzur, R. Wu, S. T albert and Y. Cho u, Journal of Engin eering for Industry 104 , 293-304 (1982).
5. P. Kazano w ski, M. E pler and W. Misiolek, Materials Scienc e and Engineeri ng A 369 , 170-180 (2 004).
6. S. Berski et al. J ournal of Materials P rocessing Technolo gy 177 , 582 -586 (2006).
7. V. Sanabria, S. Mueller and W . Reimers, Materials To day: Proce ed ings 2 , 4820 -4828 (2015).
8. P. Hor a et al., Key Engineering Materials 585 , 41 -48 (2014).
9. M. Bauser, G. Sauer and K. Siegert, S trangpressen (Aluminium -Verlag, Duesse ldorf, 200 1), pp. 265 - 266.
10. M. Mirza, D. B arton, P. Church and J. Sturges, “ Ductil fractrure o f pure cop per: an exper imental a nd numericall y
study. Journal d e Physique IV Colloq ue, 1997 , 07 (C3), pp. C3 - 891 -C3896.
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