Aluminium alloys as structural material: A review of research

Abstract

Over the last few decades aluminium alloys have been increasingly used in the construction sector due to their favourable properties. Thereafter, many research projects have been carried out with the aim to obtain a more comprehensive understanding of their structural performance and develop accurate and reliable design formulae. The scope of this paper is to provide a comprehensive review of research by discussing the reported experimental, numerical and analytical work on structural aluminium alloys. The paper presents an overview of research studies on the mechanical properties of aluminium alloys under monotonic, cyclic and thermal loading conditions. Moreover, a considerable amount of experimental and numerical investigations focussing on the structural performance and design of aluminium columns, beams and beam-columns is reviewed. The performance of connections and composite aluminium-concrete members is also discussed. Comments on the suitability of the international design specifications to structural aluminium alloys are included. Within the review, knowledge gaps are identified and the corresponding research work to fill these gaps is recommended.

Full text

1
Aluminium alloys as structural material: A review of research
Evangelia Georgantzia(a),1, Michaela Gkantou(a) , George S. Kamaris(a)
(a) Department of Civil Engineering, Liverpool John Moores University, United Kingdom
ABSTRACT
Over the last few decades aluminium alloys have been increasingly used in the construction
sector due to their favourable properties. Thereafter, many research projects have been carried
out with the aim to obtain a more comprehensive understanding of their structural performance
and develop accurate and reliable design formulae. The scope of this paper is to provide a
comprehensive review of research by discussing the reported experimental, numerical and
analytical work on structural aluminium alloys. The paper presents an overview of research
studies on the mechanical properties of aluminium alloys under monotonic, cyclic and thermal
loading conditions. Moreover, a considerable amount of experimental and numerical
investigations focussing on the structural performance and design of aluminium columns,
beams and beam-columns is reviewed. The performance of connections and composite
aluminium-concrete members is also discussed. Comments on the suitability of the
international design specifications to structural aluminium alloys are included. Within the
review, knowledge gaps are identified and the corresponding research work to fill these gaps is
recommended.
Keywords: aluminium alloys, structural response, experiments, numerical investigation,
design guidelines
1Corresponding author: Evangelia Georgantzia
Email: E.Georgantzia @2019.ljmu.ac.uk
Georgantzia E, Gkantou M, Kamaris GS. Aluminium alloys as structural material: A review
of research. Eng Struct. 2021;227:111372.

2
1 Introduction
The application of aluminium alloys as structural material has increased over the last years
owing to their favourable properties: i.e. high strength-to-weight ratio, ease of fabrication, high
degree of workability, considerable ductility, excellent thermal conductivity, high corrosion
resistance and attractive appearance at their natural finish. For this reason, 25% of the global
aluminium production is currently used in the construction sector [1]. Their ease of extrusion
makes aluminium alloys a versatile structural material allowing the production of complex
cross-sectional shapes, suitable for structures that cannot be developed from more conventional
structural materials, such as concrete or steel. Their prominent corrosion resistance makes them
well-suited for applications in marine environments without surface protection and with low
maintenance cost. Their great durability allows for structures that can maintain their inherent
properties even in large temperature variations [1]. Within the framework of sustainability and
climate-change mitigation commitments, recent technological advances led to innovative
aluminium structural systems that are more efficient from an environmental and economical
point of view compared to steel and concrete. In particular, advances on the manufacturing
process of aluminium alloys reduced the required energy more than 75% since 1995, lowering
the industry’s carbon footprint by almost 40% [2]. It has been also stated that “aluminum made
in North America is more sustainable today than ever before” [2]. Further to the decrease in
carbon dioxide emissions, structural aluminium alloys are 100% recyclable, thereby arguably
winning the title of “green metal” [3].
The aforementioned advantageous features have contributed to increased usage of aluminium
alloys in structural applications, where their application can allow for a reduction of the total
structural weight. Typical structural aluminium applications along with brief information are
presented in Figure 1. As with all structural materials, structural design codes are warranted for
aluminium alloy structures. Currently there are four international design specifications for the
structural design of aluminium alloys, as listed in Table 1.
Table 1: International Design Specifications for Aluminium Alloy Structures.
Standard ID
Standard Title [Description]
Chinese Standard: GB 50429-2007 [4]
Code for design of aluminium
structures
European Committee for Standardization:
BS EN 1999:2007 [5]
Design of aluminium structures
Australian/New Zealand Standard:
AS/NZS 1664:1997 [6]
Aluminium structures
The Aluminum Association: AA 2020 [7]
Aluminum Design Manual

3
Figure 1: Examples of aluminium alloy structures.
The Co-operative Group, Manchester, UK
The anodised exterior aluminium structure
holds the glass panels.
The Crystal, London, UK
The roof is made from 100% recycled aluminium.
Ferrari World, Abu Dhabi, UAE
The largest aluminium roof in the world. A corrugated
aluminium sheet was selected as roof material.
The Sage Gateshead, Gateshead Quays, UK
Aluminium was used to support the glazing system.
Gaylord Texan Resort & Convention Center, Grapevine,
Texas, USA
The roof is made of a glazed aluminium framework.
The Iceberg Skating Palace, Sochi, Russia
Aluminium was one of the key materials used to the construction.
2003
2004
2004
2010
2012
2012
2013
St Mary Axe, London, UK
A curved and anodised aluminium skin was used to
integrate the raking columns with the curved façade.
Casablanca Finance City Tower, Casablanca, Morocco
The modular façade elements were made out from aluminium
instead of concrete for cost reasons.
2019

4
The purpose of this paper is to provide a comprehensive review of the experimental, numerical
and analytical research work to date on the structural performance and design of aluminium
alloy structures. Upon a brief introduction in Section 1, the material properties of aluminium
alloys are discussed in Section 2. In Sections 3, 4 and 5, studies focussing on the structural
performance of columns, beams and beam-columns are presented, respectively. Reported works
on residual stresses and web crippling of aluminium sections are summarised in Sections 6 and
7. Studies on aluminium-concrete composite structures are outlined in Section 8. Reported
research on connections is presented in Section 9. Section 10 reviews experimental and
numerical works on other aluminium structural elements. Finally, concluding remarks on the
overall investigation accompanied by suggestions for future work are presented in Section 11.
2 Material properties
2.1. Overview of aluminium alloy grades
Aluminium alloys are divided into two basic categories: wrought and cast alloys. The former
comprises alloys which are melted in a furnace and then poured into moulds, whereas the latter
includes alloys treated in a solid form. Depending on the strengthening working conditions
aluminium alloys can be classified as heat-treatable and not heat-treatable. The Aluminum
Association Inc. classifies the wrought alloys into 9 series using a four-digit system and each
series comprises different combinations of alloying additions [2]. The first digit (Xxxx)
indicates the principal constituent alloy, whereas the second digit (xXxx) indicates the
modifications made in the original alloy. The last two digits (xxXX) are arbitrary numbers so
that the specific alloy can be identified in the series. Thus, the material properties can vary
offering several options for applications. Research on aluminium alloys in terms of their
structural response has focussed on wrought alloys and particularly on 5xxx and 6xxx series
that are the most attractive for structural engineering applications due to their mechanical
properties [8-10]. The alloy classification is also followed by the temper designation in order
to provide more information about the fabrication treatment. The temper designation consists
of five basic tempers; F, O, H, W, or T, accompanied by additional digits for more details about
the fabrication treatment, as described in Table 2.

5
Table 2: Summary of basic tempers for wrought alloys and the corresponding subdivisions
(adapted from [9]).
Basic tempers for wrought alloys
Subdivisions of basic tempers
F (fabricated)
The thermal conditions during working
or strain-hardening process to obtain
specific material properties do not
demand any special control.
-
O (annealed)
Treatment under high-temperature
conditions in order to achieve
maximum workability, toughness and
ductility.
-
H (strain-hardened)
Used for non-heat-treatable alloys cold
worked by strain-hardening method in
order to stabilise their strength.
The first digit indicates the type of the
thermal treatment and the second the
amount of strain-hardening.
W (solution heat
treated)
Applied to alloys subjected to natural
aging after the solution heat treatment.
Rather limited designation.
-
T (thermally
treated)
Used for heat-treatable alloys subjected
to natural or artificial aging in order
stable tempers different than F, O, or H
to be elaborated.
The first digit indicates the main type
of heat treatment and the second to
fifth [if they exist] the amount of stress
release and other special treatments.
2.2. Material properties under monotonic loading
A series of tensile coupon tests have been conducted in a wide spectrum of aluminium alloys
available in the market, aiming to investigate their material properties. Typical engineering
stress-strain curves of commonly investigated structural aluminium alloys are presented in
Figure 2 and typical mechanical properties are summarised in Table 3. In this table, E is the
Young’s Modulus, f0.2 is the stress at 0.2% strain (also known as proof stress), fu is the ultimate
stress and n is the hardening exponent according to Ramberg – Osgood constitutive model [11].
A stress-strain curve of conventional structural carbon steel [12] is also included in Figure 2 for
comparison purposes. As it can be seen in Figure 2, the stress-strain relationship of the
aluminium alloys is characterised by a rounded curve without a distinct yielding point contrary
to carbon steel. The initial material behaviour is linear elastic and is defined to relatively low
stress, f0.01, that corresponds to strain of 0.01%. After this point the material exhibits non-linear
elastic behaviour up to f0.2 stress, whilst beyond this point, plastic strains occur. Note that the
f0.2 or proof stress constitutes a threshold after which the stress-strain curve presents a “knee”
followed by a strain-hardening branch. On the other hand, carbon steel behaves similarly at the
elastic range but with larger and stiffer slope, followed by a clearly defined yield plateau and
strain-hardening branch. Comparing the stress-strain curves from different aluminium series in
Figure 2, it is apparent that 7xxx series have higher yield stress, but lower ductility compared
to 6xxx series. It can also be seen from Table 3 that more pronounced ductility is observed for
6063-T5 and 6082-T4 and more evident strain-hardening is exhibited by 6082-T4 with f0.2/fu
equal to 0.54. The yield and tensile strengths of additional commonly used structural aluminium
grades are presented for reference in Figure 3, where f0.2 and fu have been reported in the range
of 80 to 275 MPa and 160 to 350 MPa, respectively [5].

6
Figure 2: Stress-strain curves from corresponding tensile coupon tests [12-15].
Table 3: Mechanical properties of commonly investigated aluminium alloys.
Author(s) (date)
[Reference]
Aluminium
grade
f0.2
[MPa]
fu
[MPa]
E
[GPa]
f0.2/fu
n
Alsanat et al. (2019) [13]
5052-H36
211.6
257.8
64.2
0.82
-
Su et al. (2014) [14]
6061-T6
234.0
248.0
66.0
0.94
12
Su et al. (2014) [14]
6063-T5
179.0
220.0
69.0
0.81
10
Moen et al. (1999) [15]
6082-T4
120.1
221.0
66.9
0.54
26
Moen et al. (1999) [15]
6082-T6
312.2
324.2
66.7
0.96
74
Moen et al. (1999) [15]
7108-T7
314.0
333.4
66.9
0.94
65
Figure 3: Yield and tensile strengths of commonly used aluminium grades.
0
100
200
300
400
500
600
0 5 10 15 20 25
Stress [MPa]
Strain [%]
carbon steel [12]
5052-H36 [13]
6063-T5 [14]
6061-T6 [14]
6082-T4 [15]
6082-T6 [15]
7108-T7 [15]
0
100
200
300
400
500
600
700
fu [MPa]
f0.2 [MPa]

7
In order to simulate the stress-strain response of aluminium alloys, the Ramberg-Osgood model
[11] can be applied. Further to this, Baehre [16] proposed a satisfactory analytical approach,
but was unable to capture the observed “knee” of the experimental stress-strain curves. De
Matteis et al. [17] modified Baehre’s law on the basis of experimental evidence improving its
suitability. Guo et al. [18] investigated the material properties of 6061-T6 aluminium alloy and
found that the stress-strain relationship derived from the Ramberg-Osgood model [11]
combined with the Steinhardt Suggestion [19] allowed precise capture of its mechanical
behaviour. It is noteworthy that the Steinhardt Suggestion [19] greatly simplifies the description
of the constitutive relationship as it determines the hardening exponent n without considering
the 0.1% stress (f0.1). Wang et al. [20] performed a series of tensile coupon tests on 6082-T6
aluminium alloys and proposed a constitutive model based on the Ramberg-Osgood law,
combined with the application of the fast-simulated annealing method for the calculation of n.
2.3. Material properties under cyclic loading
The ductility and energy dissipation of structural materials are of great significance for the
response of structural members subjected to seismic loading. As can be seen in Table 4, there
is lack of reported works on the cyclic behaviour of aluminium alloys, which sets limitations
on their usage in earthquake prone areas. Early attempts to obtain an understanding of the
hysteretic behaviour of aluminium alloys date back to 1990s. Hopperstad et al. [21] performed
uniaxial cycling tests on specimens made from 6060 in tempers T4 and T5 under constant and
varying strain amplitudes. They suggested an amendment to the cyclic plasticity model of
Chaboche [22], so that the Bauschinger effect of temper T4 is precisely considered. Aiming to
further investigate T4 aluminium alloys, the same authors conducted biaxial proportional and
non-proportional cycling tests and extended the previous constitutive model to capture the
observed influence of the strain range and the strain path shape on the material hardening [23].
The aforementioned tests could not clarify the presence of hardening behaviour, due to the low
strain amplitudes (<2%) during the cyclic tests. To this end, Dusicka & Tinker [24] investigated
the hysteretic response of coupons generated by 6061-T6/511 alloys subjected to constant strain
amplitudes beyond 2%. The observed slight increase of the cyclic softening behaviour indicated
its potential for seismic retrofit applications. De Matteis et al. [17] conducted cyclic tests on
coupons of an almost pure aluminium alloy coded 1050A-H24 and found that it has substantial
dissipative capacity largely for higher applied strain levels. More recently, Guo et al. [25]
proposed a new constitutive model for the hysteretic behaviour of 6082-T6 and 7020-T6 on the
basis of the monotonic curve and the reduction factor method. Based on the above, more cyclic
tests are suggested to be performed to cover a wider range of aluminium alloys available in the
market.

8
Table 4: Summary of tests on material properties of aluminium alloys under cyclic loading.
(in chronological order from most recent research)
Author(s) (date)
[Reference]
Aluminium grade
Strain range [%]
Guo et al. (2018) [25]
6082-T6, 7020-T6
up to 4
Dusicka & Tinker (2013) [24]
6061-T6/511
2-4
De Matteis et al. (2012) [17]
1050A-H24
-
Hopperstand et al. (1995) [21,23]
6060-T4, 6060-T5
up to 1.2
2.4. Material properties of Heat-Affected Zone
A noteworthy characteristic of aluminium is that when high strength heat-treated aluminium
alloys (6xxx series) are welded in order to be joined with adjacent structural members, the
strength in the vicinity of the welded region is decreased significantly. This is an important
demerit of these particular aluminium alloys which cannot be neglected during the design. The
inferior material properties of this localised region around the welds, known as Heat-Affected
Zone (HAZ), are considered through the application of softening factors. According to AA 2020
[7], the HAZ extends about 25.4 mm around the weld. The influence of the HAZ on the
structural behaviour of beams and columns was demonstrated by Lai & Nethercot [26] using
numerical analysis. Mazzolani [27] determined that the parent metal strength can be reduced
almost 50% due to the presence of HAZ in 6xxx series aluminium alloys, whereas Zhu & Young
[28] found that the proof stress can undergo a decrease up to 70%.
2.5. Material properties at elevated temperatures
Since 1990s a remarkable amount of studies on the material properties of aluminium alloys
under fire conditions has been reported. Kaufman [29] significantly contributed to this research
field by conducting steady state tests on 158 different aluminium alloys and found that the
Young’s Modulus (E) is independent of the heating rate. Langhelle [30] and Hepples & Wale
[31] investigated the structural response of 6082 subjected to steady state thermal conditions.
Faggiano et al. [32] emphasised on the way that elevated temperatures affect the material
hardening factor and proposed a modified stress-strain relationship based on the Ramberg-
Osgood expression. Maljaars et al. [33] performed tests on 5083-O/H111 and 6060-T66 and
modified the Dorn-Harmathy creep model [34,35] so that to be applicable for 6xxx series
aluminium alloys. Furthermore, Kandare et al. [36] modified the Larson-Miller model [37] on
the basis of fire tests on coupons formed by 5083-H116. The reported test results were used for
the assessment of a thermo-mechanical model developed by Kandare et al. [38] as well as an
advanced modelling approach for fire conditions proposed by Feih et al. [39]. More recently,
Chen et al. [40] investigated experimentally the post-fire behaviour of 6061-T6 and 7075-T73
and suggested simplified design formulae. Su & Young [41] presented a series of empirical
equations regarding the mechanical properties of 6063-T5 and 6061-T6 exposed to fire. In the
same study, design specifications were assessed, showing that the present partial factors lead to
conservative design predictions. This is shown in Figure 4, where the test results from both
steady and transient tests appear far from the EN 1992-1-2 [42] design curve. The studies, also,
concluded that the behaviour under fire conditions is complex and dependent on the chemical

9
composition of each aluminium alloy. Additional tests that will allow more accurate design
models for each aluminium alloy ensuring both economy and safety are necessary.
Figure 4: Comparison between test results and EN 1999-1-2 [42] predictions (adapted from
[41]).
3 Columns
3.1. Local buckling
The design resistance of an aluminium structural member under compression is governed by
the cross-section classification. This is a codified procedure that implicitly treats local buckling
phenomenon, i.e. the buckling of the constituent plate elements of a cross-section under
compression. EN 1999-1-1 [5] classifies the cross-sections in four classes, using cross-section
slenderness limits (dependent on the boundary conditions of the constituent plate elements of a
cross-section), the plate element stress distribution and the heat-treatment method. Classes 1, 2
and 3 comprise cross-sections capable of yielding without failing due to local buckling, while
in Class 4 sections local buckling occurs in the elastic range and thus a reduced cross-sectional
area is considered for the evaluation of the cross-sectional resistance.
Aiming to study local buckling and the cross-sectional performance, early tests on stub columns
have been reported [43-47]. More recently, a considerable amount of stub column tests have
been conducted in a wide range of cross-sectional shapes (Figure 5), aluminium grades and
width-to-thickness ratios of the most slender constituent plate element. Zhu et al. [48]
investigated the behaviour of plain and lipped channel (C-) stub columns, whereas Mazzolani
et al. [49] tested angles and proposed an empirical equation about the local buckling resistance.
Liu et al. [50,51] studied the local buckling behaviour of stiffened and irregular-shaped cross-
sections and Yuan et al. [52] evaluated experimentally the post-buckling behaviour of slender
(i.e. large width to thickness ratio) I-sections. Wang et al. [53] conducted stub columns tests on
CHSs made from 6082-T6, whilst Feng & Young [54] dealt with perforated cross-sections.
Following, Feng et al. [55,56] determined the reduced load-bearing capacity due to the presence
0.0
0.2
0.4
0.6
0.8
1.0
1.2
0 100 200 300 400 500 600
fy,T/fy
Temperature, T [°C]
EN 1999-1-2 [42]
steady state tests
transient state tests

10
of holes by testing perforated stub columns with rectangular hollow sections (RHSs), square
hollow sections (SHSs) and circular hollow sections (CHSs). Upon experimental testing on
tubular sections, Su et al. [14] highlighted the significant contribution of the material strain-
hardening on the cross-section capacity and assessed the applicability of the Continuous
Strength Method (CSM) [57,58], that was originally developed for stainless steel stocky (i.e.
small width to thickness ratio) cross-sections. Su et al. [59,60] extended the CSM to cover
aluminium sections and proposed new slenderness limits as well as an effective thickness
formula on the basis of collected data.
The studies are listed in Table 5, where the design code assessment is also shown. The mean
values and coefficient of variation (COV) of ratios of the reported test strength, Nu, to the code
predicted strength, Npred, are included. For mean ratios Nu/Npred higher than unity, the
predictions are conservative, for lower than unity they are unsafe and for close to unity they are
accurate. Furthermore, high values of COV suggest scattering and thus the predictions are
considered as unreliable. As it can be seen, excessively conservative predictions were reported
for channel sections in [48], which is opposed to an economic design process. In general, only
a few studies indicated accurate cross-sectional strength predictions. The lack of accuracy is
also related to the fact that the design formulae for aluminium often adopt similar principles to
structural steel design, without sufficient consideration of the differences between the two
materials. Modifications in line with obtained test data on aluminium are needed.
Figure 5: Cross-sectional shapes employed in stub column investigations.
(a) rectangular hollow
section (RHS)
(b) square hollow
section (SHS)
(c) circular hollow
section (CHS)
(d) I-section
(e) unequal angle (L-)
(f) equal angle (L-)
(g) plain channel (C-)
(h) lipped channel (C-)
(i) RHS with internal cross
stiffeners
(j) SHS with internal cross stiffeners
(k) irregular shape

11
Table 5: Summary of aluminium alloy stub column tests.
(in chronological order from most recent research)
Author(s)
(date)
[Reference]
Aluminium
grade
Shape
No of
tests
Width-
to-
thickness
ratios
Design codes
Nu/Npred
Assessment
mean
COV
Zhu et al.
(2019) [48]
6063-T5,
6061-T6
plain C-,
lipped C-
8
25.50-
25.90
GB 50429-2007 [4]
1.65
0.06
conservative
EN 1999-1-1:2007 [5]
1.35
0.07
conservative
AS/NZS 1664.1:1997 [6]
1.28
0.15
conservative
AA [7]
1.28
0.15
conservative
NAS [125]
1.21
0.10
conservative
CSM [59,60]
1.12
0.16
conservative
Feng et al.
(2018) [55]
6063-T5,
6061-T6
perforated
RHS, SHS
16
27.30-
43.67
NAS [125]
0.92
0.11
unsafe
Feng et al.
(2016) [56]
6063-T5,
6061-T6
perforated
CHS
10
23.48-
49.81
NAS [125]
1.50
0.11
conservative
Wang et al.
(2015) [53]
6082-T6
CHS
9
14.00-
26.70
-
Feng &
Young
(2015) [54]
6061-T6
perforated
SHS
28
6.20-
48.30
AISI 2008 [121]
0.96
0.32
unsafe
NAS [125]
0.95
0.33
unsafe
Yuan et al.
(2015) [52]
6061-T6,
6063-T5
I-
15
35.70-
71.70
GB 50429-2007 [4]
1.13
0.12
conservative
EN 1999-1-1:2007 [5]
1.12
0.12
conservative
AS/NZS 1664.1:1997 [6]
1.10
0.09
conservative
AA [7]
1.06
0.09
conservative
Liu et al.
(2015) [51]
6063-T5
irregular
7
-
-
Liu et al.
(2015) [50]
6063-T5
stiffened
closed-
sections
10
-
GB 50429-2007 [4]
0.96
0.05
accurate
EN 1999-1-1:2007 [5]
1.01
0.04
accurate
AA [7]
0.94
0.08
unsafe
DSM [65]
0.83
0.04
unsafe
AISI 2008 [121]
0.98
0.08
accurate
Su et al.
(2014) [14]
6061-T6,
6063-T5
SHS, RHS
(with and
without
internal
stiffeners)
15
3.20-
20.70
EN 1999-1-1:2007 [5]
1.07
0.09
conservative
AS/NZS 1664.1:1997 [6]
1.34
0.16
conservative
AA [7]
1.19
0.16
conservative
CSM [59,60]
1.04
0.06
accurate
Mazzolani
et al. (2011)
[49]
6xxx
angles
64
2.90-
35.40*
-
*calculated according to available data.
3.2. Flexural buckling
The flexural buckling behaviour of aluminium alloy columns has been under thorough
investigation, being one of the primary constituents for the assurance of the structural integrity.
According to the current design guidelines, buckling classes are determined by two material
groups based on the temper designation, as shown in Table 6.

12
Table 6: Material groups based on temper designation. (adapted from Wang et al. [53])
Specifications
Material group 1
Material group 2
GB 50429-2007 [4]
T6
Other tempers
EN 1999-1-1:2007 [5]
T6, H14/24/34
Other tempers
AS/NZS 1664.1:1997 [6]
T5, T6, T7, T8, T9
O, H, T1, T2, T3, T4
AA 2020 [7]
T5, T6, T7, T8, T9
O, H, T1, T2, T3, T4
Note: Material groups 1 and 2 refer to buckling curves A and B in EC9, respectively.
In order to comprehend the ultimate performance of aluminium alloy columns, early studies
have been reported by Hopperstad et al. [45] who tested 6082-T4 and 6082-T6 columns, and
Manevich [61] who numerically investigated the influence of the material strain-hardening on
the critical buckling stress. Over the last years, a considerable amount of experimental studies
has been performed, as summarised in Table 7. In this table, the test boundary conditions and
the slenderness ratio, Le/r, of the specimens are also included (Le is the effective buckling length
and r the radius of gyration of the cross-section). Wang et al. [53] focussed on the reliability
level of current design rules on CHSs columns, while Adeoti et al. [62] expanded the
investigation on columns formed by H-sections and RHSs. Wang et al. [63] studied L-shaped
columns manufactured by 7A04 high-strength aluminium alloy, whereas Wang et al. [64]
focussed on I-section columns. Feng et al. [56] investigated the buckling behaviour of
perforated columns, suggesting that a properly modified Direct Strength Method (DSM) [65],
a design approach suggested for cold-formed steel sections, could be suitable for the design of
CHS columns with circular openings. After two years, Feng et al. [55] reported that the DSM
cannot be applied for the design of perforated RHS and SHS columns. The aforementioned test
results were also used by Feng & Liu [66] to conduct an extensive parametric study and adjust
the EN 1999-1-1 [5] equations, taking into account the reduced cross-sectional area due to
perforation. A numerical study on irregular-shaped sections was carried out by Chang et al. [67]
who concluded that the DSM is able to predict the interactive buckling failure mode accurately
but not in every case. Recently, Wang et al. [68] tested columns with large RHS and I- sections
and Zhu et al. [48,69] presented their test results on plain and lipped channel columns. As shown
in Table 7, the reported test data have been used to assess current design rules and it can be
concluded that the international guidelines are overly conservative and confirm the need for
further research into this field. In addition, many of past studies have focussed on hollow
sections, which are less prone to torsional failure. Hence, despite the exhaustive experimental
and numerical investigation on the structural response of columns, test data on interactive
torsional-flexural buckling behaviour are relatively limited and further research is
recommended.

13
Table 7: Summary of aluminium alloy column experiments.
(in chronological order from most recent research)
Author(s)
(date)
[Reference]
Aluminiu
m grade
Shape
No of
tests
Boun
dary
condit
ions
Slende
rness
ratio
[Le/r]
Design codes
Nu/Npred
Assessment
mea
n
COV
Zhu et al.
(2019) [48]
6063-T5,
6061-T6
plain C-,
lipped C-
20
fixed
ends
-
GB 50429-2007 [4]
1.38
0.20
conservative
EN 1999-1-1:2007 [5]
1.45
0.14
conservative
AS/NZS 1664.1:1997 [6]
1.23
0.16
conservative
AA [7]
1.23
0.16
conservative
NAS [125]
1.21
0.15
conservative
Wang et al.
(2018) [68]
6061-T6
I-, RHS
7
pinned
ends
28.96-
116.74
GB 50429-2007 [4]
1.55
0.25
conservative
EN 1999-1-1:2007 [5]
1.30
0.22
conservative
AA [7]
1.06
0.19
conservative
Feng et al.
(2018) [55]
6063-T5,
6061-T6
perforated
RHS, SHS
21
pinned
ends
13.94-
93.22
NAS [125]
0.97
0.07
accurate
Wang et al.
(2017) [64]
6063-T5,
6061-T6
I-
11
fixed-
pinned
ends
46.90-
67.50
GB 50429-2007 [4]
1.45
0.13
conservative
EN 1999-1-1:2007 [5]
1.45
0.11
conservative
AS/NZS 1664.1:1997 [6]
1.27
0.09
conservative
AA [7]
1.13
0.13
conservative
Feng et al.
(2016) [56]
6063-T5,
6061-T6
perforated
CHS
8
pinned
ends
28.84-
58.88
NAS [125]
1.27
0.12
conservative
Wang et al.
(2016) [63]
7A04
L-
42
pinned
ends
15.00-
100.00
GB 50429-2007 [4]
2.76
0.27
conservative
EN 1999-1-1:2007 [5]
1.21
0.21
conservative
AA [7]
1.19
0.34
conservative
Adeoti et al.
(2015) [62]
6082-T6
H-, RHS
30
pinned
ends
22.36-
163.01
GB 50429-2007 [4]
1.14
0.09
conservative
EN 1999-1-1:2007 [5]
1.14
0.09
conservative
AA [7]
1.20
0.09
conservative
GB 50017-2003 [103]
1.21
0.10
conservative
Wang et al.
(2015) [53]
6082-T6
CHS
15
pinned
ends
24.42-
73.99*
EN 1999-1-1:2007 [5]
1.10
0.08
conservative
AS/NZS 1664.1:1997 [6]
0.97
0.13
accurate
AA [7]
1.14
0.13
conservative
*calculated according to available data.
3.3. Welded columns
As mentioned in Section 2.4, the reduced strength of the HAZ affects the structural response of
the structural member and thus it should be considered during the design process. To this end,
Zhu & Young [28,70-74] examined the buckling behaviour of RHSs, SHSs and CHSs columns
with and without transverse welds. They proposed new design criteria for the ultimate strength
based on the DSM and new values for HAZ softening factors. Zhu et al. [75] extended this
investigation to channel sections and modified the DSM and the CSM approach to make them
applicable to welded channel columns. Feng et al. [55] dealt with perforated RHS and SHS
columns incorporating welded and non-welded specimens. Their experimental outcomes
demonstrated the applicability of the design criteria proposed by Zhu & Young [72] to welded
columns.
3.4. Columns at elevated temperatures
In order to comprehend the buckling response and design of columns at elevated temperatures,
experimental and numerical work has been performed, as listed in Tables 8 and 9, respectively.
Langhelle & Amdahl [76] performed column buckling tests to clarify the consequences of the
viscoplastic behaviour at elevated temperatures. Suzuki et al. [77] conducted a series of column

14
tests under fire conditions and extended the simple plastic theory to estimate the critical
temperature beyond which column failure occurs. Maljaars et al. [78,79] carried out axial
compression tests and finite element (FE) analyses on slender SHSs and angles under steady
and transient state conditions and proposed new less conservative cross-section classification
limits for EN 1999-1-2 [42]. In a following numerical work, Maljaars et al. [80,81] pointed out
that the stress-strain relationships at elevated temperatures are more curved than at ambient
temperature and that the buckling resistance is directly linked to the inelastic critical stress. Liu
et al. [82] determined the buckling behaviour of columns with irregular-shaped cross sections
by numerical means and suggested a modification to the equations provided by EN 1999-1-2
[42]. In a more recent study, Jiang et al. [83] performed tests and FE models on RHS and CHS
columns and modified the stability coefficient of EN 1999-1-1 [5] and GB 50429 [4] to take
into account the effect of the elevated temperatures on the normalised slenderness and the
imperfection parameter.
Table 8: Summary of tests on columns at elevated temperatures.
(in chronological order from most recent research)
Author(s)
(date)
[Reference]
Type of test
Aluminium
grade
Shape
No of
tests
Temperature
[°C]
Design codes
Assessment
Jiang et al.
(2018) [83]
axial
compression
6061-T66
RHS,
CHS
108
up to 400
-
Maljaars et al.
(2009) [78]
axial
compression
5083-H11,
6060-T66
SHS, L-
55
up to 330
EN 1999-1-2
[42]
conservative
Suzuki et al.
(2005) [77]
fire resistance
test [loaded
and non-
loaded
heating]
5083-O,
5083-H112
box, H-
23
up to 850
-
Langhelle &
Amdahl
(2001) [76]
axial
compression
6082
-
31
-
-

15
Table 9: Summary of numerical investigations on columns at elevated temperatures.
(in chronological order from most recent research)
Author(s)
(date)
[Reference]
Type of test
Aluminium
grade
Shape
No of
analyses
Temperature
[°C]
Design codes
Assessment
Jiang et al.
(2018) [83]
axial
compression
6063-T5,
6061-T6,
6063-T6,
6061-T4
RHS, CHS,
J-, T-, L-,
C-, Z-,
T- [one
sym. axis]
8829
up to 400°C
-
Liu et al.
(2016) [82]
axial
compression
6061-T6
irregular
shaped
300
up to 500°C
GB 50429-2007 [4]
conservative
EN 1999-1-1:2007 [5]
conservative
EN 1999-1-2 [42]
conservative
AA [7]
conservative
DSM [65]
conservative
AISI 2008 [121]
conservative
Maljaars et al.
(2009)
[79,80,81]
axial
compression
5083-O/H111,
6060-T66
SHS, I-
48
200,300
EN 1999-1-2 [42]
conservative
4 Beams
4.1. Flexural resistance
The flexural resistance and rotational capacity of beams are of significant importance in order
to ensure the safe transfer of the vertical loads to the foundation. This is one of the earliest
research topics, since the first experimental works date back to 1950s, when Panlilo [84]
investigated the behaviour of two-span statically indeterminate beams. Later, Mazzolani et al.
[85] extended the plastic design to aluminium alloy structures and Welo [86] performed tests
under uniform moment and determined the moment-curvature behaviour. Thereon, numerous
experimental and numerical investigations have been carried out on aluminium beams under 3-
, 4- and 5-point bending conditions, as summarised in Table 10. Opheim [87] conducted 4-point
bending tests and found that there is no significant difference between tensile and compressive
behaviour of 6060-T4 beams. Moen et al. [15,88] demonstrated through experimental and
numerical studies that the rotational capacity is dependent on the material strain-hardening and
the magnitude of the moment gradient. Their test results [15] were used by De Matteis et al.
[89] who proposed new limits on the cross-section classification of EN 1999-1-1 [5],
considering the material strain-hardening. The importance of the material strain-hardening was
also highlighted by Su et al. [90-93]. In another study, Zhu & Young [94] modified the current
DSM achieving more accurate and reliable design provisions for flexural SHS members. Kim
& Peköz [95] developed a new formulation for the stress at ultimate limit state based on test
results of doubly symmetric I-section beams. Kim & Peköz [96] also presented a simplified
design approach named Numerical Slenderness Approach in order to determine the nominal
stresses of each constituent plate element of a complex section under flexure. The reliability of
the proposed method was evaluated by performing a series of tests on beams with mullion
sections. Castaldo et al. [97] numerically studied the ultimate behaviour of RHS beams under
non-uniform bending and proposed multivariate non-linear equations for their ultimate flexural

16
resistance and rotational capacity. Piluso et al. [98] extended the aforementioned study to I-
sections fabricated by 6082-T4 and 6063-T5. Experimental and numerical studies on perforated
CHS beams subjected to gradient and constant moments were reported by Feng et al. [99,100].
They found that the presence of holes, their size and number reduce the flexural capacity.
Recently, Montuori et al. [101] reported a thorough finite element investigation on I-beams
formed by high-yielding low-hardening aluminium alloys. The outcomes denoted that the
increased values of slenderness parameter and shear length ratio reduce the rotational capacity.
Focussing on lateral-torsional buckling, Cheng et al. [102] investigated numerically the lateral
stability of I-section beams and suggested a modification to the GB 50017-2003 [103]. The
proposed modified design methodology was assessed by Wang et al. [104] concluding that it
provides more accurate predictions compared to EN 1999-1-1 [5]. A few years later, Wang et
al. [105] extended their investigation conducting experiments on I-beams including specimens
with and without intermediate stiffeners subjected to concentrated loads.
Table 10 summarises the studies and the design code assessment by providing the mean and
COV values of the reported obtained ultimate flexural strengths (Mu) over design strengths
predicted by the international design codes (Mpred). The overall high Mu/Mpred ratios reveal
largely conservative design estimations. The latter can also be visualised in Figure 6, which
presents reported Mu values normalised by Mpred of EN 1999-1-1 [5], and plotted against the
cross-sectional slenderness parameter (b/t, i.e. width to thickness ratio). In addition, as shown
in Table 10, there are only a few reported studies on 5-point bending tests and hence additional
experiments are suggested to better evaluate the plastic performance of indeterminate beams.
Table 10: Summary of investigations on beams.
(in chronological order from most recent research)
Author(s)
(date)
[Reference]
Type of
study
Aluminium
grade
Type of
bending
test
Shape
No
of
tests
Design codes
Mu/Mpred
Assessment
mean
COV
Montuori et al.
(2020) [101]
FE
6061-T6,
6082-T6
3-point
H-, I-
240
-
Feng et al.
(2020) [100]
Exp
6061-T6,
6063-T5
3-point,
4-point
perforated
CHS
8
NAS [125]
1.20
0.23
conservative
Feng et al.
(2019) [99]
FE
6061-T6,
6063-T5
3-point,
4-point
perforated
CHS
408
-
Piluso et al.
(2019) [98]
FE
6082-T4,
6063 T5
3-point
H-, I-
240
-
Kim & Peköz
(2018) [96]
Exp &
FE
6063-T5
4-point
mullion
2 &
-
AA [7]
conservative
Castaldo et
al.(2017) [97]
FE
6082-T6
3-point
RHS
252
-
Wang et al.
[105]
Exp &
FE
6061-T6,
6063-T5
simply
supported
I-
10 &
24
EN 1999-1-
1:2007 [5]
1.40
0.10
conservative
Su et al. (2016)
[93]
Exp &
FE
6063-T5,
6063-T6
3-point,
4-point,
5-point
SHS,
RHS with
internal
stiffeners
30 &
150
EN 1999-1-
1:2007 [5]
1.41
0.11
conservative
AS/NZS
1664.1:1997 [6]
2.11
0.21
conservative
AA [7]
1.67
0.18
conservative
CSM [59,60]
1.30
0.10
conservative

17
Su et al. (2015)
[91,92]
Exp&
FE
6061-T6,
6063-T5
5-point
SHS,
RHS
27&
120
EN 1999-1-
1:2007 [5]
1.82
0.23
conservative
AS/NZS
1664.1:1997 [6]
2.26
0.23
conservative
AA [7]
2.02
0.26
conservative
CSM [59,60]
1.39
0.16
conservative
Su et al. (2014)
[90]
Exp &
FE
6061-T6,
6063-T5
3-point,
4-point
SHS,
RHS
29 &
132
EN 1999-1-
1:2007 [5]
1.17
0.11
conservative
AS/NZS
1664.1:1997 [6]
1.54
0.16
conservative
AA [7]
1.38
0.14
conservative
CSM [59,60]
1.11
0.11
accurate
Wang et al.
(2012) [104]
Exp
6061-T6,
6063-T5
simply
supported
I-
40
EN 1999-1-
1:2007 [5]
0.92
0.13
unsafe
Kim & Peköz
(2010) [95]
Exp &
FE
6063-T6
4-point
I-
3 &
-
AA [7]
1.21
0.06
conservative
Zhu & Young
(2009) [94]
Exp &
FE
6061-T6,
6063-T5
4-point
SHS
10 &
60
EN 1999-1-
1:2007 [5]
1.31
0.11
conservative
AS/NZS
1664.1:1997 [6]
1.38
0.20
conservative
AA [7]
1.35
0.20
conservative
DSM [65]
1.21
0.07
conservative
Cheng et al.
(2006) [102]
FE
-
simply
supported
I-
250
-
De Matteis et
al. (2001) [89]
FE
6082-T4,
6082-T6
4-point
RHS
-
EN 1999-1-1 [5]
conservative
Moen et al.
(1999) [88]
FE
6082-T4,
6082-T6,
7108-T7
4-point
unwelded
I-, welded
I-, box
19
-
Moen et al.
(1999) [15]
FE
6082-T4,
6082-T6,
7108-T7
4-point
unwelded
I-, welded
I-, box
38
EN 1999-1-1 [5]
1.15
0.11
conservative
Opheim (1996)
[87]
Exp &
FE
6060-T4,
6064-T6
4-point
SHS
-
-

18
Figure 6: Comparison between test results and design predictions using EN 1999-1-1[5].
4.2. Welded beams
Until now relatively little attention has been given at the behaviour of aluminium alloy welded
beams. Thus, the existing design approaches adopt the same principles to the corresponding
ones for steel welded beams, leading to gross approximations, since the two materials differ
considerably. Moen et al. [15,88] and Matusiak [106] presented their studies on welded I-beams
highlighting the significant reduction on rotational capacity due to welding. The reported results
by the latter were used by Wang et al. [107] who focussed on the vicinity of the weld and
determined its impact on the total strength and ductility of the beams. The authors proposed a
new modelling methodology for the region around the weld, assuming shell elements,
geometric imperfections, plastic anisotropy, inhomogeneous material properties and ductile
failure. The actual structural performance of welded beams is still ambiguous and thus more
experimental research needs to be carried out.
4.3. Beams at elevated temperatures
A few studies have been carried out on the structural behaviour of beams exposed to fire. Suzuki
et al. [77] performed fire resistance tests and proposed fire design formulae, whereas more
recently, Meulen et al. [108] performed 3-point bending steady and transient state tests up to
300°C in order to assess the EN 1999-1-2 [42]. The experimental and numerical studies on
aluminium beams at elevated temperatures are limited and thus more 3-, 4- and 5- point bending
tests on different cross-sectional shapes in a wide range of applied strain rates are necessary.
Data from these experiments will form a database that can be used for the development of more
accurate design models and more reliable design provisions for aluminium alloys.
0
0.5
1
1.5
2
2.5
3
0 10 20 30 40 50
Mu/Mpred
b/t
Moen et al. [15]
Su et al. [90]
Su et al. [91]
Su et al. [93]
Zhu & Young [94]

19
5 Beam-columns
The behaviour of aluminium alloy members under combined compression and bending has also
been reported. Clark [109], Klöppel & Bärsch [110] and Gilson & Cescotto [111] performed
tests on RHS, I- and T stocky sections. Zhu & Young [112,113] presented their experimental
work on SHS, RHS and CHS specimens under combined axial compression and bending about
the weak axis. The obtained outcomes demonstrated that the predicted beam-column strengths
are underestimated by the current design guidelines. Zhao et al. [114-116] reviewed the Chinese
Standards and proposed new values for certain design parameters. Furthermore, Zhu et al. [117]
tested eccentrically compressed I-shaped members under elevated temperatures and presented
a simplified correlation curve able to predict the bearing capacity of columns subjected to
eccentric compression up to 300°C.
6 Residual stresses
The residual stresses developed during the manufacturing process can have a significant impact
on the overall structural response of a member. According to Mazzolani [10], the residual
stresses are quite low for extruded profiles, heat treated or not, and thus can be ignored, contrary
to welded profiles, where the residual stresses can have a significant impact on the load-bearing
capacity of the structure. Aiming for a better understanding of the residual stress distribution in
aluminium sections, Huynh et al. [118] investigated the residual stresses of cold-rolled
aluminium channel sections using the sectioning method. It was shown that the in-plane residual
stresses were significant only in the corner parts of the smallest and thinnest C-section, whereas
the out-of-plane residual stresses were considerable (up to 30% of yield stress) for all the
investigated sections. Similar findings for cold-formed steel open sections have been reported
by Moen et al. [119] and Gardner & Cruise [120]. There is a need to extend this investigation
in various cold-formed, hot-rolled and welded cross-sections, so that the effect of residual
stresses will be adequately considered in the design process.
7 Web crippling
Web crippling is specified as localised buckling and yielding of the web in the vicinity of the
applied concentrated load. Research works examining the web crippling of a plethora of
sections including end-two-flange (ETF), interior-two-flange (ITF), end-one-flange (EOF) and
interior-one-flange (IOF) loading and boundary conditions, as defined in AISI 2008 [121], have
been reported and summarised in Table 11. The first reported work was presented by Tryland
et al. [122] who found that the web thickness and the flange stiffness considerably affects the
ultimate capacity. Later Zhou & Young [123,124] conducted an extensive investigation on
SHSs and RHSs in a wide slenderness range and suggested modified design formulae to the
North American Specification (NAS) [125]. Zhou et al. [126] tested SHSs under concentrated
bearing loads and proposed threshold slenderness values beyond which the web buckling
becomes the predominant failure mode. Zhou & Young [127] extended their investigation on
SHSs with perforated webs proposing a strength reduction factor and a new web crippling
design equation for SHSs with circular web holes. Chen et al. [128] studied further the web

20
crippling behaviour of SHSs proposing new equations for the ultimate capacity. In another
study, Su & Young [129] proposed a more accurate and reliable design methodology for the
web bearing capacity of stocky sections which takes into account the significant effect of the
material strain-hardening. Alsanat et al. [13,130] tested for first time roll-formed aluminium
lipped channel sections under ETF and ITF conditions and proposed modified rules on the basis
of the obtained test data. Recently, Zhou & Young [131] carried out tests on plain and lipped
channel sections with restrained flanges. The assessment of the design specifications based on
the most crucial loading-boundary condition are also summarised in Table 11, revealing the
current lack of accuracy and reliability in the design predictions of the web crippling
phenomenon. Contrary to the cross-sectional (Table 5), column (Table 9) and flexural strengths
(Table 10) that are generally underestimated by the codes, Table 11 shows that overall the
codified capacities against web crippling are not safely estimated.
Table 11: Summary of web crippling tests.
(in chronological order from most recent research)
Author(s)
(date)
[Reference]
Aluminium
grade
Shape
No
of
tests
Loading-
boundary
conditions
Web
slenderness
ratios (b/t)
Design codes
Nu/Npred
Assessment
mean
COV
Zhou & Young
(2020) [131]
6063-T5,
6061-T6
lipped C-,
plain C-
52
ETF, ITF
43.00-58.00
EN 1999-1-
1:2007 [5]
0.75
0.26 (ETF)
unsafe
1.18
0.32 (ITF)
AS/NZS
1664.1:1997 [6]
1.00
0.47 (ETF)
conservative
1.06
0.28 (ITF)
AA [7]
1.00
0.47 (ETF)
conservative
1.06
0.28 (ITF)
NAS [125]
1.12
0.36 (ETF)
unsafe
0.62
0.40 (ITF)
Alsanat et al.
(2019) [13]
5052-H36
lipped C-
40
ETF, ITF
3.33-10.00*
AS/NZS
1664.1:1997 [6]
0.50
0.37 (ETF)
unsafe
0.88
0.24 (ITF)
EN 1993-1-
3:2005 [132]
0.49
0.06 (ETF)
unsafe
0.60
0.07 (ITF)
Su & Young
(2018) [129]
6063-T5,
6061-T6
SHS,
RHS
34
ETF, ITF,
EOF, IOF
2.80-28.00
EN 1999-1-
1:2007 [5]
0.53
0.23 (EOF)
unsafe
0.78
0.19 (IOF)
0.87
0.14 (ETF)
1.16
0.11 (ITF)
AS/NZS
1664.1:1997 [6]
0.62
0.32 (EOF)
unsafe
0.78
0.30 (IOF)
0.96
0.14 (ETF)
1.00
0.16 (ITF)
AA [7]
0.47
0.20 (EOF)
unsafe
0.78
0.30 (IOF)
0.96
0.14 (ETF)
1.00
0.16 (ITF)
EN 1993-1-
3:2005 [132]
3.14
0.31 (EOF)
conservative
1.01
0.21 (IOF)
5.05
0.30 (ETF)
8.04
0.24 (ITF)
AISC [133]
0.54
0.25 (EOF)
unsafe
0.88
0.22 (IOF)
0.87
0.12 (ETF)
1.19
0.12 (ITF)
-
SHS
48
30.00-88.00
2.45
0.53 (EOF)
conservative

21
Chen et al.
(2015) [128]
ETF, ITF,
EOF, IOF
EN 1993-1-
3:2005 [132]
1.47
0.33 (IOF)
2.26
0.52 (ETF)
1.34
0.33 (IOF)
GB 50017 [103]
0.29
0.54 (EOF)
unsafe
0.42
0.39 (IOF)
0.28
0.53 (ETF)
0.38
0.44 (IOF)
Zhou & Young
(2010) [127]
6061-T6
perforated
SHS
84
ETF, ITF
6.20-49.50
EN 1999-1-
1:2007 [5]
0.75
0.23 (ETF)
unsafe
0.95
0.15 (ITF)
AA [7]
0.95
0.47 (ETF)
accurate
0.97
0.29 (ITF)
Zhou et al.
(2009) [126]
6061-T6
SHS
64
ETF, ITF
6.20-48.30
EN 1999-1-
1:2007 [5]
1.04
0.25 (EL)
accurate
1.05
0.20 (IL)
AA [7]
1.86
0.37 (EL)
conservative
1.46
0.25 (IL)
Zhou & Young
(2008) [123]
6063-T5,
6061-T6
SHS,
RHS
150
EF**, IF***
6.30-74.50
-
Tryland et al.
(1999) [122]
6082-T6
SHS, I-
52
-
-
-
*calculated according to available data.
**EF: End-bearing Loading
***IL: Interior-bearing Loading

22
8 Composite structures
8.1. Aluminium-concrete structural members
Following similar concept and principles with the composite steel-concrete structures and in
particular with the concrete-filled steel tubes (CFST), the possibility of combining aluminium
with concrete has been investigated. Research work on the structural response of concrete-filled
aluminium tubes (CFAT) with typical cross-sections shown in Figure 7 has been reported.
(a) RHS
(b) SHS
(c) CHS
(d) double-skin CHS
(e) CHS reinforced with CFRP
Figure 7: Investigated cross-sections of aluminium-concrete composite members [134-142].
Zhou & Young [134] conducted axial compression tests on concrete-filled aluminium stub
columns with SHSs and RHSs and concluded that the AS/NZS 1664.1:1997 [6] and AA [7]
design codes are generally unconservative. Later, Zhou & Young [135,136] extended their
experimental investigation on CHS stub columns filled with concrete and developed design
criteria considering the observed material interaction. In a more recent study, Zhou & Young
[137] assessed experimentally the compressive response of concrete-filled double-skin tubes
and suggested formulae for their ultimate capacity. Wang et al. [138] used the data reported by
Zhou & Young [135] and evaluated whether the “nominal yield strength” method adopted by
GB 50936 [139] for CFST is applicable to CFAT, concluding that it provides conservative but
reliable predictions.
Feng et al. [140] tested simply-supported concrete-filled SHS and RHS beams, whereas Chen
et al. [141] performed 4-point bending tests on concrete-filled CHS beams. In both
investigations, the ultimate strength almost doubled thanks to the concrete infill, which
prevented premature failure due to local buckling. Chen et al. [142] investigated the flexural
behaviour of concrete-filled CHSs strengthened by carbon fibre-reinforced polymer (CFRP). It
was observed that the slightly improved ultimate capacity was accompanied by a reduction in

23
the ductility. Modifications on the Architectural Institute of Japan (AIJ) standards [143] so as
to consider the contribution of the CFRP reinforcement were also presented.
More research on this field should be carried out in order to adequately determine the structural
behaviour of CFATs and propose design criteria able to achieve efficient exploitation of both
materials. Future studies could include flexural buckling tests on CFATs with and without
CFRP strengthening, beam-column tests, stub columns under eccentric compression and
investigation of their behaviour at elevated temperatures.
8.2. Aluminium-CFRP structural members
Wu et al. [144] were at the forefront of strengthening aluminium alloy tubular sections against
web crippling using CFRP, finding that the web crippling capacity can experience almost a
four-fold increase due to the CFRP. Islam & Young [145] focussed on the effect of the
application of six different types of adhesives and fibre-reinforced polymers (FRPs) on the web
crippling capacity of SHSs and RHSs. It was shown that the higher the web slenderness ratio,
the greater the enhancement of the web crippling strength. This was confirmed at their
following experimental work [146] and the reported results were used in a recent numerical
study where the NAS [125] design equations were modified in order to consider the contribution
of both the CFRP-strengthening and the adhesive to the web crippling capacity [147].
9 Connections and joints
9.1. Welded
Owing to the difficulties related to the weldability of aluminium [148], only limited work has
been reported to date on aluminium welded connections. Early attempts for a comprehensive
understanding of the behaviour of welded connections were made by Soetens [149], who
investigated experimentally and numerically the structural response of welded connections in
RHSs fabricated by 6063-T5 and 7020-T6. His findings were incorporated in the international
specifications for the design of aluminium alloy structures (ECCS [150], NEN 3854 [151], CP
118 [152]). Another research study on welded connections was performed by Chan & Porter
Goff [153] who evaluated experimentally the effects of the reduced strength zone on the
ultimate capacity, ductility and failure mode of 7xxx series aluminium alloys. The structural
response of welded T-stub joints under monotonic tensile loading was examined by De Matteis
et al. [154] and it was shown that EN 1999-1-1 [5] equations provide reliable although slightly
underestimated design predictions. The scarcity of the reported data reveals the need of
additional experiments on aluminium welded connections to enable a better understanding of
their behaviour.

24
9.2. Bolted
Over the last two decades, a series of studies have been performed on bolted connections under
various arrangements and load cases, as illustrated in Figure 8. Table 12 summarises the
reported experimental work. De Matteis et al. [155,156] conducted a thorough experimental
and numerical work on T-stub connections under monotonic and cyclic loading. Kim [157]
carried out tests on single shear bolted connections and found that the curling effect (out-of-
plane deformation) reduces suddenly the ultimate capacity. These findings were used by Cho
& Kim [158] who modified the strength equations for block shear fracture and bearing factor,
taking into account the curling effect. In a more recent study by Wang et al. [159], twenty bolted
connections were tested under tensile loading and the obtained results were used for the
assessment of GB 50429 [4], EN 1999-1-1 [5] and AA [7] design codes, concluding that the
aforementioned design specifications lead to conservative predictions. De Matteis et al. [160]
carried out an extensive parametric study on the structural behaviour of T-stub joints showing
that the material strain-hardening and the ductility considerably affect the strength of the joint.
Brando et al. [161] determined the ultimate capacity of the web in beam-to-column joints
subjected to tension and adjusted the design criteria developed for steel joints by using
correction factors that consider the mechanical characteristics of aluminium alloys. Recently,
Adeoti et al. [162] reported a study dealing with the flexural behaviour of hexagonal bolted
joints underling the importance of considering all the parameters with great impact on the
structural behaviour and stiffness in order to design joints with high performance.
Figure 8: Configuration of investigated bolted connections.
(a) T-stub bolted connection under
monotonic and cyclic loading (adapted from
De Matteis et al. [155,156])
(b) Bolted connection under tensile loading
(adapted from Wang et al. [159])
bolts
bolts
aluminium
alloy plate
steel plates
steel cover
plates
alloy plate
T-stub
head
nut
thread
bolt

25
Table 12: Summary of aluminium alloy bolted connection/joint tests.
(in chronological order from most recent research)
Author(s) (date)
[Reference]
Aluminium
grade
Connection/joint
type
No
of
tests
Design
codes
Nu/Npred
Assessment
mean
COV
Adeoti et al. (2019) [162]
6082-T6
hexagonal bolted
joints
6
-
Wang et al. (2018) [159]
6061-T6, 6063-
T5
shear connection
in single shear
with two bolts
20
EN
1999-1-
1:2007
[5]
1.36
0.03
conservative
AA [7]
1.42
0.11
conservative
GB
50429-
2007 [4]
2.78
0.12
conservative
Kim (2012) [157]
6061-T6
shear connection
in single shear
with four bolts
10
-
De Matteis et al. (2004)
[155]
6061-T6, 6082-
T6, 7020-T6
welded plates
with holes
26
-
Over the last five years, there is also a wide usage of aluminium alloy gusset (AAG) pinned,
rigid or semi-rigid joints in practice. Guo et al. [163,164] performed a series of tests on fourteen
AAG joints in order to define their out-of-plane flexural response. The results were used to
elaborate simplified design formulae about the resistance against block tearing and local
buckling. Guo et al. [165] adapted the component method included in EN 1993-1-8 [166] for
AAG steel joints system and proposed suitable expressions for their bending behaviour. Guo et
al. [167] investigated the flexural response of AAG joints exposed up to 300°C and proposed
design criteria for the bearing capacity and the non-linear flexural stiffness. In a further study
by Guo et al. [168], the hysteretic behaviour of AAG joints was assessed through cycling
loading tests. Shi et al. [169] conducted experiments on two-way AAG joints subjected to pure
bending and shear loading and they proposed a theoretical model able to accurately capture the
mechanical behaviour of these joint systems. Liu et al. [170] determined experimentally the
flexural behaviour of double- and single- layer AAG joints. Comparisons between the two types
of the investigated joints demonstrated the superior structural response of the former.
Additional research in order to obtain a better understanding of the structural response of bolted
connections under various configurations, loading cases (static, cyclic and fatigue) and
aluminium alloy types, is recommended. This will allow for design criteria able to take into
account this complex behaviour and lead to safe and economic design solutions.
10 Other studies
Kesawan et al. [171] conducted experimental work on the flexural response of mullions caused
by wind pressure and suction, whereas the following year a numerical study on long span
mullions with complex-shape sections under wind suction was presented by them [172].
Scheperboer et al. [173] studied numerically the buckling behaviour of perforated steel and

26
aluminium plates and suggested that the design rules for steel perforated plates are applicable
to aluminium alloy plates. Pursuing optimised cross-sectional shape with efficient exploitation
of the material distribution, Tsavdaridis et al. [174] applied Structural Topology Optimization
in aluminium cross-sections. They concluded that further research should be conducted
including more global and local failure modes. Ampatzis et al. [175] suggested a useful
methodology for determining the safety factor of spatial aluminium frame structures against
elastoplastic collapse. He et al. [176] proposed a novel modular support structure assembled by
a foldable plane frame and joints suitable for temporary structures. Finally, the hysteresis
behaviour of aluminium shear panels has been investigated, demonstrating their potential as
dissipative devices in seismic resistant structures [177-179]. Related to this, it is noteworthy
that studies on the seismic behaviour of columns and beams remain scarce. Therefore, a series
of tests on structural members subjected to cyclic loading would be an interesting future
research field in terms of the investigation of their ductility and energy dissipation capacity.
11 Conclusions and future work
This paper reviewed the reported research work on structural aluminium alloys, providing a
complete view of their mechanical properties, structural response and design of basic structural
elements. The history of structural aluminium’s investigation is relatively short and thus more
research is needed in order to obtain a thorough understanding of its behaviour. On the basis of
the reviewed papers, the following conclusions can be drawn:
1. Overall the current design guidelines do not provide accurate strength predictions,
which are opposed to an economical and efficient design philosophy. This is related to
the fact that their formulae are based on limited amount of experimental and numerical
results. Design codes sometimes adopt similar principles to their steel structure
counterparts, without sufficient consideration of the differences between the two
materials.
2. Despite the advantageous features of structural aluminium alloys members, the
investigation revealed that there are still limitations in their design, forcing the designers
to favour more conventional materials.
3. Topics with limited number of studies that have been mentioned throughout this work
are summarised in Table 13 as future recommendations. Additional research work can
lead to modifications of the existing design codes and potentially increase structural
engineers’ confidence towards a more frequent employment of aluminium alloys.
4. Finally, scope of future work is to bridge the gap between theoretical and real world,
making aluminium alloy an alternative construction material, capable of efficiently
responding to the challenges encountered in real-life structures.

27
Table 13: Summary of recommended future work.
Investigation topic
Methods of investigation (experimental & numerical)
Material properties
under cyclic loading
Cyclic tests coupons in a wide range of aluminium alloys.
Material properties at
elevated temperatures
Coupon tests of various alloys under fire conditions in order to
develop more accurate design models considering the chemical
composition.
Interactive torsional-
flexural buckling of
columns
Column tests on open cross-sections of various aluminium alloys.
Flexural response
Bending tests on different cross-sectional shapes of welded beams.
Inelastic performance of statically indeterminate beams
Flexural response at
elevated temperatures
Bending tests including various cross-sectional shapes in a wide
range of applied strain rates.
Influence of residual
stresses on the structural
performance
Measurements of magnitude and distribution of residual stresses in
sections of various fabrication processes.
Structural response of
concrete-aluminium
elements
Structural members under various loading scenarios at room and
elevated temperatures.
Structural response of
connections
Welded and bolted connection tests under various configurations,
loading cases (static, cyclic, fatigue) and aluminium alloy types.
Seismic behaviour of
columns and beams
Cyclic tests to investigate ductility and energy dissipation capacity.
Acknowledgements
The financial support provided by Liverpool John Moores University for the Doctoral Program
of Higher Education is greatly acknowledged.
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