design of composite columns made of concrete ... - CCVI Information

DESIGN OF COMPOSITE COLUMNS MADE OF CONCRETE FILLED
TUBES WITH INNER MASSIVE CORE PROFILES AND HIGH
STRENGTH MATERIALS
Univ.-Prof. Dr.-Ing. G. Hanswille
Institute for Steel and Composite Structures, University of Wuppertal
Wuppertal, Germany
stahlbau@uni-wuppertal.de
Dipl.-Ing. M. Lippes
Gevelsberg, Germany
lippes@email.de
ABSTRACT
This paper deals with the design of composite columns made of concrete filled tubes with inner
massive core profiles as well as high strength steel and high strength concrete. New research
work is presented regarding the effect of residual stresses in inner core profiles resulting from
the cooling process during fabrication. Based on Finite Element studies new models are
presented for the determination of the distribution of residual stresses in solid round steel bars
taking into account the steel grade, the dimensions of the core profiles and the cooling
conditions during fabrication. Additional theoretical investigations are presented to estimate the
effects of creep and shrinkage under long term loading. Finally recommendations are given for
the design of this type of composite column based on the methods given in Eurocode 4-1-1.
1 INTRODUCTION
The use of composite columns made of high strength steel and high strength concrete is
beneficial in order to realise a high resistance in combination with small dimensions of cross
sections. At present the design of composite columns according to Eurocode 4 Part 1-1 /1/ is
applicable only for steel grades up to S460 and concrete strengths up to C50/60 because no
sufficient test data are available for higher strength. In the following new test results and
analytical investigations are presented to open the scope in Eurocode 4-1-1 especially for high
strength concrete. To fulfil the demand for more slender columns the use of concrete filled tubes
with additional massive inner core profiles is very
effective. Due to limited information about the effects of
residual stresses in massive core profiles which
develop during production, the use of massive round
steel cores in general is also not yet covered by the
standards. Therefore numerical investigations for the
determination of residual stresses in massive round
steel bars had been made in order to widen the scope
in EN 1994-1-1 /1/ for this type of column.
Fig. 1- Composite column with inner
massive core profile

2 NEW TESTS AND EXPERIMENTAL SETUP
The test program comprises 6 columns with concrete filled tubes (series C1) and further 6
columns with additional massive inner core profiles with a diameter of 70 mm (series C2).
Columns in axial compression and with uniaxial bending due to end eccentricity e o of 5mm,
20 mm and 60 mm were performed. Information about the test specimen and the test program
is given in Figure 2 and Table 1. The concrete cylinder strength was 95 N/mm 2 , the yield
strength of the tubes 333 N/mm 2 and the yield strength of the massive round core was 411
N/mm 2 in the centre of the profile and 447 N/mm 2 at the surface.
test series C1
tube 323,9 x 4
F
e o
crushing of concrete
test series C2
240
plate
480x360x80
3792 mm
longitudinal cracks
local buckling
core ∅ 70
e o
100
centring bar
360
cracks in the
tension zone
80
r=30
e o
480
Fig. 2 – Test setup
Fig. 3 – typical failure mode
According to EC4-1-1 appendix B the columns were preloaded. First the load was applied in
increments up to approximately 50% of the expected failure load and then cycled 150 times
between 25% and 50% of the expected failure load. During the cyclic loading the load was
applied force controlled. After this procedure the load was increased and applied displacement
controlled up to the failure load. Each test was held at specific load levels for more than 15
minutes to get information about relaxation. In this manner the ultimate static load considering
relaxation effects could be determined. In all tests a nearly constant reduction of approximately
4% of the ultimate short time static load could be observed. When reaching the ultimate load all
tests failed in the middle third of the column by crushing of concrete and local buckling of the
steel tube (Figure 3).
Two columns did not reach the expected failure load. The unexpected failure was caused by
airlocks due to an inefficient compaction of the concrete in the middle of the column and under
the end plates. The concrete had been filled from the top. Based on the results it can be stated
that this method and a rapid settling of the concrete leads to the risk of an insufficient
compaction of the concrete especially when using high strength concrete. Because of the
unexpected results these columns were excluded from further analyses. The horizontal
deflection under maximum loads and the strains measured in the middle of the columns were
used to determine the internal forces and the effective load eccentricity at failure load. Using the
experimental stress-strain curves the bending moment in the middle of the column was
calculated with the assumption that the cross section remains plain. The evaluation of the

tangential strains in the tubes indicated that the confinement effect of the tube on the
compressive strength of concrete can be neglected.
The evaluation of the internal forces showed that the effective eccentricity e x for the
recalculation of the columns depends on the load level and is given by e x = e o +Δe x according to
Figure 4. At the beginning of loading the effective eccentricity was approximately equal to the
distance between the axis of the load introduction and the centre line of the centring bar. By
increasing the load and also the rotation at the bearing of the column the eccentricity moved to
the edge of the centring bar. The evaluation showed that the effective eccentricity converges to
a value Δe x,u with increasing loading so that the load eccentricity at failure could be exactly
determined (Figure 4).
Table 1 – Experimental and calculated ultimate loads and corresponding horizontal
deformations
test series C1
test series C2
test 1 3 4 5 6 1 2 3 4 6
experimental values
e 0 [mm] 20 60 60 5 5 20 20 60 60 5
F u [kN] 5439 3192 2986 5980 5903 5560 5322 3392 3277 7359
Δe x,u [mm] 10 14 15 13 15 10 17 10 11 7
u x (F u ) [mm] 12.4 20.2 19.6 15.1 11.6 16.0 16.1 23.8 23.2 15.0
e x,u =e o +Δe xu +u x 42.4 94.2 94.6 33.1 30.6 46.0 53.1 93.8 94.2 27.0
results of the recalculation with the FE - model
e o,cal [mm] 30 74 75 18 20 30 37 70 71 12
F u [kN] 4981 3247 3188 6182 5938 5493 5094 3401 3409 6969
u x (F u ) [mm] 17.5 26.5 25.4 17.4 17.9 20.5 22.2 28.4 29.6 14.
e x,u =e o,cal +u x 47.5 100.5 100.4 35.4 37.9 50.5 59.2 98.4 100.6 26.3
Δe x [mm]
F
F
40
30
20
e x
u x
10
0,2
0,4
0,6
0,8
Δe x,u F
1,0 F u
Δe x
e o
e x
Fig. 4 – Relation between the effective eccentricity and the load level

3 NUMERICAL INVESTIGATIONS AND EVALUATION OF TEST RESULTS
For the recalculation of the columns a 3-dimensional model with cuboid elements of the FEprogram
Ansys was used. The mechanical properties of steel were represented by a multi-linear
elastic-plastic stress-strain relationship on the basis of the von Mises yield criterion. For the
mechanical properties of concrete an additional condition was used to model the low resistance
against tension stresses more accurately /3/. In order to take into account the bond behaviour
between steel and concrete in the interfaces between steel and concrete contact elements were
used with a friction coefficient µ = 0.5. The model allows to consider initial bow imperfections
and additional eccentric positions of the massive inner steel core. A typical relation between the
normal force and the bending moment at midspan is shown in Figure 6. The measured and
calculated strains in Figure 6 demonstrate that for design it can be assumed that the cross
section remains plain.
In order to check the safety level in comparison with the simplified design methods in Eurocode
4 the general design method in Eurocode 4 has to be used. Regarding the safety concept in
case of non linear finite element analysis the Eurocodes for composite structures give no clear
guidance. Therefore the method developed in /4, 5/ was used, where for the resistance an
overall safety factor γ R has to be determined which depends on the relation between the acting
normal force and the acting bending moment. The safety factor γ R can be determined using the
full plastic cross-section interaction curve of the section. As shown in Figure 5, the interaction
curve using the mean values (measured values in case of tests) for the material strength as well
as the curve using the design values according to Eurocode 4 have to be calculated. For the
determination of the interaction curve based on design values the partial factor 1.5 for concrete,
1.1 for structural steel and 1.15 for reinforcement have to be used. For a given combination of
internal forces N Ed and M Ed the safety factor γ R is given by the ratio of the vectors R pl,m and R pl,d .
The incremental finite element analysis has to be performed using an initial geometrical bow
imperfection and considering residual stresses in the massive inner core profile. It has to be
verified, that the load amplification factor λ u is greater than the overall safety factor γ R for the
resistance.
σ c
f cm
σ s
0,4 f c
E cm
ε c1
ε c1u
N Ed
N
f ct
E d
R pl,d
R pl,m
B
M Ed
concrete
A
ε c
f sm
M
f tm
E s
reinforcement
σ a
E v
Rpl,m
γR
=
Rpl,d
- ε s
F d
F
w d
F u,m
= λ u
F d
w u
f u
+
f y
w
-
E a
ε v
geometrical
imperfection
F
e o
w o =L/1000
structural steel
w o
ε a
residual
stresses in the
core profile
σ E
-
d K
d K
d
Fig. 5 – Safety concept in case of non linear design of composite columns /4,5/

The test results were evaluated based on the statistical method given in EN 1990, Appendix D
in order to determine characteristic and design values. This method is based on a comparison
of the experimental F u,exp and theoretical F u,cal values of the ultimate load, where the theoretical
values result from a mechanical model. For the columns the theoretical values F u,cal were
determined by help of a finite element analysis taking into account the measured stress strain
relations from the tests and considering residual stress in the core profiles determined according
to the following chapter 4. For the calculation of the ultimate load F um in addition to the end
eccentricities acc. to Table 1 geometrical bow imperfections with a parabolic shape and a
maximum value L/1000 were assumed. The results of the statistical evaluation of the test results
acc. to EN 1990 are shown in Figure 7. The characteristic values are given by F R,k = 0.9 F u,m
and the appropriate design value result to F R,d = 0.79 F u,m . For comparison in Figure 7 also the
design values resulting from the general design method in Eurocode 4-1-1 are shown. These
values calculated with the partial factors acc. to Eurocode 4-1-1 and the safety concept are
shown in Figure 5. It can be seen that the general design method in Eurocode 4-1-1 leads to
safe results also in case of high strength materials.
F [kN]
Test C2-6
Strains at midspan
F u,exp , F d [MN]
8000
6000
4000
FEM
ε [‰] strain gauges
F 2
F 1
-4
FE-model
F
-3
-2
F 2 = 6969 kN
-1
F 1 = 6000 kN
8,0
6,0
4,0
characteristic
values F Rk
design values F Rd
mean values F um
2000
Test
100
200
M [kNm]
N
M
strain
gauges
2,0
statistical evaluation acc.
to EN 1990
design values based on
the general design
method in EN 1994-1-1
2,0
4,0
6,0
8,0
F u,cal
Fig. 6 - Relation between the normal force and the
bending moment at midspan
and strain distribution
Fig. 7 – Comparison of the experimental
and theoretical ultimate loads and test
evaluation acc. to EN 1990
4 RESIDUAL STRESSES IN MASSIVE ROUND STEEL BARS
As mentioned before for the determination of ultimate loads of columns, structural and
geometrical imperfections must be taken into account. For composite columns with inner
massive core profiles especially the residual stresses in the steel cores have to be considered.
Residual stresses arise during production due to the uneven temperature distribution while the
steel bar is cooling off. The surface areas are cooling down and hardening faster with ongoing
cooling. The later cooling of the inner parts causes then compression in the already hardened
part near the surface and tension stresses in the inner parts of the cross section. In composite
columns these residual stresses cause a reduction of flexural stiffness and a reduction of the
ultimate load due to earlier yielding of structural steel and increasing yielding zones in the
structural steel section.

The simplified design method in Eurocode 4-1-1 covers the influence of residual stresses by
equivalent initial bow imperfections where the distribution of residual stresses in structural steel
sections is taken into account by different buckling curves. For columns with round inner core
profiles until now no information is given in Eurocode 4-1-1 regarding the relevant buckling
curve and the corresponding member imperfections. Up to now only a few experimental
investigations are known which give information about the distribution of residual stresses in
massive round steel bars /6/.
To obtain more detailed information about the stress distribution and the maximum stresses, the
cooling process was simulated using the FE-program ANSYS. The residual stresses were
calculated for different diameters and steel grades. The temperature dependent material
properties were taken from DIN EN 1993-1-2 /7/. Furthermore it was assumed, that there is no
subsequent treatment to diminish the residual stresses.
First of all the time-dependent temperature distribution was determined. The calculations were
performed for the period of cooling after the final rolling of the steel. Thereby the convective and
radiative heat transmission coefficients and the initial temperature of the massive profile were
the most important parameters which influence the distribution and values of the residual
stresses. The simulations were performed with two different cooling conditions. Additional to
normal cooling conditions unfavourable environmental conditions were examined. To achieve
this, the convective heat transfer coefficient was determined for different continuous air flow. For
the normal cooling conditions an air flow rate of w=1 m/s was assumed. With an air flow rate of
w=5 m/s extremely unfavourable cooling conditions were taken into account. This extreme
cooling condition was used, to get information about the relation between the residual stresses
and the different cooling conditions. The interaction between convection and radiation was not
considered because this leads to a more slowly cooling and therefore to smaller residual
stresses. Considering the interval of cooling from the initial state down to 100 °C the
environmental conditions lead to cooling durations as shown in Figure 8. Figure 9 shows the
development of the temperature in the centre and at the surface for a solid bar with a diameter
of 200 mm.
ΔT [K]
T [°C]
T R
T i
T a
w
300
1200
ΔT = T i –T R.
24
16
12
6
d K
t (T R = 100 o C) [h]
Case 1:
T i = 1000°C,
T a = 20°C and
α c (w = 5 m/s)
Case 2:
T i = 1200°C,
T a = 20°C and
α c (w = 1 m/s)
d k [mm]
200 400 600
Fig. 8 –Duration of cooling down to a
temperature T R =100 o C at the surface
of the solid profile
250
200
150
100
50
1000
800
600
400
200
temperature T i
temperature
difference ΔT
0 1 2 3 4 5 6 7
T i
d k =200mm
T R
normal cooling conditions
extreme cooling conditions
t [h]
Fig. 9 – Time dependent development of the
temperature and the temperature difference for a solid
bar with a diameter d k = 200 mm

With the temperature dependent stress-strain relationships according to /7/ the distributions of
the residual stresses can be determined from the calculated time-dependent temperature
distributions. Different environmental conditions affect the residual stresses significantly. For
normal cooling conditions the distribution is characterised by nearly constant tensile stresses in
the inner part of the cross section. When the cooling speed is increased the stress distribution
can be approximated by a parabolic function. The area with constant tensile stresses in the
centre becomes more distinctive for higher steel grades (Figure 10).
σ E (r) [N/mm 2 ]
200
100
0
-100
-200
S235
S460
approximation
S355
r
r K
0,2 0,4 0,6 0,8 1,0
d k = 200 mm
-400
-300
-200
-100
σ ED [N/mm 2 ]
extreme cooling
conditions
S460
approximation
S355
S235
normal cooling
conditions
σ ED -
+
r
r
r K
d K
Approximation:
dK
2
σED
= σED,o
[N/ mm ]
200
2
σED,o
= −125
N/ mm
⎡ 2
2r ⎤
σ E (r) =σ ED ⎢1
−
2 ⎥
⎢⎣
r K ⎥⎦
200 400 600
d k [mm]
Fig. 10 – Influence of the steel
grade on the distribution of residual
stresses in solid round bars
Fig. 11 – Maximum compressive residual stresses at the
surface of the cross section as a function of the steel
grade and the diameter d k
Figure 11 shows that the approximation for maximum residual compressive stress at the surface
of solid bars and the distribution of the residual stresses is in good agreement with the more
accurate values. Because the residual stresses near the surface are mainly responsible for the
reduction of the ultimate loads for the recalculation of the tests the approximation was used.
5 INFLUENCES OF CREEP AND SHRINKAGE
Under permanent loads creep and shrinkage of concrete in composite members cause a
redistribution of the sectional internal forces from the concrete section to the structural steel
section. This can lead in case of further loading to an earlier yielding in the steel section, to
larger yielding zones and therefore to a significant reduction of the flexural stiffness at ultimate
limit states. For columns in compression and bending the decrease of the flexural stiffness
causes increasing horizontal deformations and additional second order effects. Eurocode 4-1-1
takes into account these effects by a reduction of the flexural stiffness of the column. The tests
were subjected to short-term loading only. To obtain information about the influence of creep
and shrinkage more exact calculations were done, where the creep function for high strength
concrete according to Eurocode 2-1 /8/ and the test results in /9/ were used. The tests with
concrete filled hollow sections in /9/ show that for concrete filled tubes a reduced creep
coefficient can be assumed, which is in the range of 25% of the value for a cross section with its

surface fully exposed to the ambient environmental conditions. In a first step the redistribution of
sectional forces and the horizontal deformations were determined for permanent loads by an
incremental analysis taking into account second order effects. In a second step the ultimate
loads were determined by a non-linear calculation taking into account the redistribution of
internal forces and the additional deformations caused by creep and shrinkage due to
permanent loads and the effects of residual stresses, geometrical imperfections and non-linear
material behaviour. The calculations were carried out for a nominally pinned column in axial
compression with a representative cross-section with a concrete filled steel tube 406.4 x 8.8,
steel grade S355 and concrete strengths of 60 N/mm² and 100 N/mm². Table 2 shows the
results of four columns with a related slenderness of 1.0 and 2.0 where F Rd,c+s is the design
value of the ultimate load taking into account creep and shrinkage and F Rd is the ultimate load
without creep and shrinkage effects. It can be seen that in case of concrete filled tubes with high
strength concrete the effects of creep and shrinkage lead to a reduction of the ultimate load less
than 5% even if the column has a high related slenderness. Comparative results can be
achieved by using the simplified design method given in EC4-1-1.
Table 2 – Effects of creep and shrinkage and comparison with the simplified design method in
EN 1994-1-1
compressive
cylinder strength
f c
related
slenderness
λ k
f c = 60 N/mm 2 1.0
f c = 60 N/mm 2 2.0
f c = 100 N/mm 2 1.0
f c = 100 N/mm 2 2.0
creep
coefficient
ϕ(t ∞ ,t 0 )
Exact non-linear
calculation
F Rd
[kN]
F +
L/w o
Rd,c s
FRd
Simplified method acc. to
Eurocode 4-1-1 with equivalent
imperfections w o acc. to chapter 6
F Rd
[kN]
0 5912 5679
0.97 360
0.25x 1.34 5722
5555
0 2112 2037
0.98 440
0.25x 1.34 2077
1941
0 7425 7437
0.97 360
0.25x 0.873 7224
7324
0 2869 2842
0.98 440
0.25x 0.873 2844
2724
F Rd,c + s
FRd
0.98
0.95
0.98
0.96
6 PROPOSALS FOR A SIMPLIFIED DESIGN METHOD BASED ON EUROCODE 4
The use of the general design method in Eurocode 4 based on a geometric and physical
nonlinear calculation can not be recommended for practical design because of an enormous
calculation effort. In the following a simplified method is presented which is based on the
simplified design method in Eurocode 4-1-1 and shown in Figure 12. Internal forces have to be
determined by second order linear elastic analysis with an effective flexural stiffness (EI) eff,II .
This stiffness accounts for the cracking of concrete and the yield zones in the structural steel
section (factors K o and K e,II in Fig. 12). The influence of geometrical and structural imperfections
is taken into account by equivalent geometrical bow imperfections, which depend on the type of
cross-section. In case of axial compression and bending it has to be verified that in the critical
cross-section the design value of the second order bending moment M Ed does not exceed the
design value of bending resistance M Rd of the cross-section. The bending resistance results
from the plastic M-N interaction curve of the cross-section and an additional correction factor α M
which takes into account the difference between the plastic and the non-linear bending
resistance with strain limitations.

N pl,Rd
N
Non-linear resistance
with strain limitation
F Ed
=N Ed
μ M pl,Rd
plastic
interaction curve
N Ed
μα M
M pl,Rd
L
w o
M
plastic distribution of stresses
M pl,Rd
f yd,R
f yd,K f cd
M pl,Rd,N
-
-
-
+
N Ed
stresses in case of strain limitation in concrete
(EJ) eff
Effective flexural stiffness
( EJ) eff = Ko
(EaJa
+ Ke,II
Ecm
Jc
)
Verification
ε
f yd,R
f yd,K σ c
-
-
-
α M
M pl,Rd,N
N Ed
αM
MEd
μ Mpl,Rd
≤ 1,0
Fig. 12 - Simplified design method according to Eurocode 4-1-1
1,0
0,8
0,6
0,4
N
N pl,Rd
N pl,Rd tube 406,4x8,8 S235
tube 406,4x8,8 S235 1,0
core d k
= 200 mm S235
Concrete C30/37 0,8
Concrete C30/37
0,6
0,4
N
0,2
1,0
0,8
0,6
0,2
M
M
0,2 0,4 0,6 0,8 1,0 1,2 M pl,Rd 0,2 0,4 0,6 0,8 1,0 1,2 M pl,Rd
N
N
N pl,Rd
N pl,Rd tube 406,4x8,8 S460
tube 406,4x8,8 S460 1,0
core d k
= 200 mm S460
Concrete C100/115 0,8
Concrete C100/115
0,6
0,4
0,2
with strain
limitation
plastic
0,2 0,4 0,6 0,8 1,0 1,2
M
M pl,Rd
0,4
0,2
M
M
0,2 0,4 0,6 0,8 1,0 1,2 pl,Rd
Fig.13 – Plastic and non-linear M-N-interaction curve for two typical cross sections
For high strength materials and concrete filled steel tubes with additional inner massive steel
cores the interaction curve depends strongly on the strength of the used materials. The
proportion between the plastic interaction curve and the interaction curve with strain limitation
for the concrete can differ significantly. Figure 13 shows this for typical cross-sections.
Therefore the constant values for the correction factor α M in Eurocode 4 can not be used for
columns with high strength concrete. The analysis of various cross-sections leads to the
following approximation for the correction factor α M .

N
α
Ed
M = αMo
− αN
(1)
Npl,Rd
where the values α Mo and α N are given in table 3, N Ed is the design value of the normal force
and N pl,Rd the plastic resistance to compression. Interim values can be determined by linear
interpolation.
Table 3 – values α Mo and α N
steel grade f yd,core concrete f cd d K /d= 0 d K /d= 0,75
of the core section [N/mm²] grade [N/mm²] α Mo α N α M0 α N
C30/37 20 0,90 0,10 0,85 0,15
S235 218 C 60/75 40 0,90 0,25 0,80 0,15
C100/115 60 0,90 0,40 0,75 0,15
C30/37 20 0,85 0,25 0,70 0,20
S460 418 C 60/75 40 0,85 0,35 0,60 0,20
C100/115 60 0,85 0,45 0,50 0,20
Based on the correction factor α M the equivalent bow imperfection can be determined by
comparison of the ultimate loads of the simplified method with the ultimate loads determined by
an exact non-linear calculation taking into account a geometrical imperfection of L/1000 and
structural imperfections according to chapter 4. The equivalent imperfections depend on the
cross section geometry and on the slenderness as well as on the effective flexural stiffness /4,
5/. The analysis of several columns covering different material strengths, buckling lengths,
eccentricities and cross-sections gave the following equation (2) for the maximum value of
equivalent bow imperfection w o where the factors k i are given in Table 4. The factor k 1 takes into
account the influence of the steel grade, the factor k 2 the influence of the diameter d k of the
inner core cross-section and the factor k 3 the influence of the related slenderness.
L
= 400 ⋅ k1
⋅ k 2 ⋅ k3
(2)
w 0
For two typical columns Figure 9 shows a comparison of ultimate loads determined with the
general design method according to Eurocode 4-1-1 and the ultimate loads determined with the
simplified design method according to Eurocode 4-1-1 but using the correction factor α M
according to equation (1) and the equivalent bow imperfection acc. to equation (2). The
comparison demonstrates that the proposed design method gives an excellent agreement with
the exact general design method.
Table 4 – Geometrical bow imperfection for concrete filled tubes with additional
inner core profiles
cr
steel grade
d k
d diameter d k of core
Npl,k
λ K =
related slenderness
N
S355 k 1 = 1
S460 k 1 = 1,25
≤ 200 mm
d [mm]
k 2 = 1+
K
400
> 200 mm
d [mm]
k 2 = 2 − K
400
λ 0,5 k 3 = 0,8
K ≤
λ K > 0,5 k 3 = 0,7
+ 0,2 λK

N Rd
[MN]
N Rd
[MN]
simplified design method
80,0
60,0
40,0
20,0
tube 660x7,1
d k
= 200 and 400 mm
S355 und S460
C60/75
40,0
30,0
20,0
10,0
tube 406,4x8,8
d k
= 90,110 and 300 mm
S355 und S460
C60/75
N Rd,A
[MN]
20,0 40,0 60,0 80,0
10,0 20,0 30,0
N Rd,A
[MN]
.
general design method
Fig. 9 – Comparison of the ultimate loads determined with the general design method and the
simplified design method acc. to Eurocode 4-1-1
40,0
7 REFERENCES
[1] EN 1994-1-1: Eurocode 4: Design of composite steel and concrete structures
Part 1-1: General rules and rules for buildings (2004)
[2] EN 1990: Eurocode - Basis of structural design (2002)
[3] Porsch, M: Modellierung von Schädigungsmechanismen zur Beurteilung der
Lebensdauer von Verbundkonstruktionen aus Stahl und Beton, Dissertation,
Fachbereich Bauingenieurwesen, Institut für Konstruktiven Ingenieurbau, Lehrstuhl für
Stahlbau und Verbundkonstruktionen, Bergische Universität Wuppertal (2008)
[4] Hanswille G.: Die Bemessung von Stahlverbundstützen nach nationalen und EU-Regeln,
Der Prüfingenieur 22, (2003)
[5] Bergmann, R.: Hanswille, G.: New design method for composite columns including High
strength steel , ASCE, Proceedings of the 5th international conference Composite
Construction in Steel and Concrete V, South Africa, 2004
[6] Roik, K., Schaumann, P.: Tragverhalten von Vollprofilstützen – Fließgrenzenverteilung
an Vollprofilquerschnitten, Institut für Konstruktiven Ingenieurbau, Ruhr-Universität
Bochum, Juli 1980
[7] EN 1993-1-2:2005: Eurocode 3: Design of steel structures, Part 1-2: General rules,
Structural fire design (2005)
[8] EN 1992 -1-1: Eurocode 2: Design of concrete structures, Part 1-1: General rules and
rules for Buildings (2004)
[9] Ichinose, L.H., Watanabe, E, Nakai, H.: An experimental study on creep of concrete
filled steel pipes, Journal of Constructional Steel Research 57 (2001), S. 453 – 466