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Load-Carrying Capacity of Timber–Concrete Joints with Dowel-Type Fasteners
Article in Journal of Structural Engineering · May 2007
DOI: 10.1061/(ASCE)0733-9445(2007)133:5(720)
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Load-Carrying Capacity of Timber–Concrete Joints
with Dowel-Type Fasteners
A. M. P. G. Dias 1 ; S. M. R. Lopes 2 ; J. W. G Van de Kuilen 3 ; and H. M. P. Cruz 4
Abstract: This paper analyzes the load-carrying capacity of timber–concrete joints made with dowel-type fasteners. The scope of the
work was to obtain information to be used in the design of timber–concrete composite structures. In order to achieve that, two different
approaches were used: experiments with timber–concrete joints loaded in shear, and the use of nonlinear analytical models. Shear tests
were performed using various timber species, concrete mixtures, and fasteners. The results from these tests are presented and compared
with the results determined with three different analytical models. These models differ in the approach used to simulate concrete:
linear–elastic, linear–elastic with crushing, and elastic–perfectly plastic. The analysis of the test results shows that the load-carrying
capacity of this type of joint is significantly influenced not only by the strength of the materials but also by the shape of the fastener. From
the three models proposed, the one assuming elastic–perfectly plastic behavior for concrete leads to the results closer from the test results,
however, the best statistical correlations between model results and test results were obtained with the model assuming linear–elastic
behavior with crushing.
DOI: 10.1061/ASCE0733-94452007133:5720
CE Database subject headings: Bearing capacity; Joints; Concrete; Wood.
Introduction
Dowel-type fasteners are one of the most popular fasteners for
creating timber–concrete structures, not only because they are
easy to use and have good mechanical performance, but also because
they are relatively cheap and available everywhere. Nevertheless,
its mechanical properties, for example its load-carrying
capacity and slip modulus, are not well known yet, despite being
essential in the analysis of the composite structures Van der
Linden 1999; Dias 2005. Typically, the mechanical properties of
this type of joint are evaluated either by means of experiments or
by means of analytical or numerical models.
Experimental research in many cases is still considered the
best solution to obtain relevant mechanical properties for a particular
joint configuration. Ahmadi and Saka 1993 tested seven
types of high-strength nails available to the local construction
industry Gulf Persian to connect timber imported from Asia to
ready mix concrete. Three different penetration depths were studied
on the timber side, eight, 11 and 15 times the diameter of the
fastener, and for each configuration three specimens were tested.
1 Assistant Professor, Dept. of Civil Engineering, Univ. of Coimbra
FCTUC, Polo II 3030-290 Coimbra, Portugal.
2 Associate Professor, Dept. of Civil Engineering, Univ. of Coimbra
FCTUC, Polo II 3030-290 Coimbra, Portugal.
3 Associate Professor, Dept. of Civil Engineering, TU Delft,
Stevinweg 1 2628 CN Delft, The Netherlands.
4 Senior Research Officer, LNEC National Laboratory of Civil
Engineering, Av. do Brasil, 101 1700-066, Lisboa, Portugal.
Note. Associate Editor: J. Daniel Dolan. Discussion open until
October 1, 2007. Separate discussions must be submitted for individual
papers. To extend the closing date by one month, a written request must
be filed with the ASCE Managing Editor. The manuscript for this paper
was submitted for review and possible publication on August 10, 2005;
approved on June 21, 2006. This paper is part of the Journal of Structural
Engineering, Vol. 133, No. 5, May 1, 2007. ©ASCE, ISSN 0733-
9445/2007/5-720–727/$25.00.
The objective of the work was to choose the best type of nail for
local application and the most adequate penetration depth of the
fastener. Gutkowski and Chen 1996 reported tests on four types
of nails: one double head nail and three common round nails with
diameters of 2.9, 3.3 and 3.8 mm. Each nail type was tested with
two penetration depths and, for each configuration, three tests
were performed with a concrete age of 14 days and another six
with 28 days. Dias 1999 tested timber–concrete joints made
with square nails with and without an interlayer, performing nine
tests for each configuration. Gelfi and Giuriani 1999 tested dowels
12 and 16 mm, in each case three penetrations were used:
three, four, and six times the diameter. In the tests using 16 mm
dowels, there was an interlayer formwork between the timber
and concrete. Most of this research was performed with fasteners
and materials locally available and not completely described in
most of the cases. For that reason comparison between the test
results remains difficult. Besides, it is difficult to extrapolate these
results to new situations. In the last research work Gelfi and
Guiriani 1999, a simple dowel possible to reproduce everywhere
was used, however, the number of tests was very low and the
parameters studied were too few to allow a broad analysis.
In terms of models, the load-carrying capacity of timber–
concrete joints is usually obtained using traditional models for
timber–timber joints. This option is elected for two main reasons:
similarities between mechanical behavior of timber–concrete and
timber–timber joints, and the much larger amount of test data
available from past research. In terms of timber–timber joints, a
wide review of the methods developed was made by Patton-
Mallory et al. 1997. The different methodologies and design
philosophies were described and its advantages and limitations
analyzed. CEN Eurocode 5 CEN 2003, for example, proposes
calculation of load-carrying capacity of timber–concrete joints
based on models for the timber–timber joints, but uses modification
factors. The method was evaluated by Dias et al. 2003, who
compared the model results with the results obtained in laboratory
720 / JOURNAL OF STRUCTURAL ENGINEERING © ASCE / MAY 2007

tests using dowel-type fasteners. It was concluded that the model
results were satisfactory but could be improved if the particular
specificities of timber–concrete joints, for example, the concrete
compression strength and foundation modulus, were taken into
consideration.
Gelfi et al. 2002 proposed a formulation specifically for
timber–concrete joints. It proposes calculation of the loadcarrying
capacity using a plastic limit analysis considering one
timber member and one concrete member, both materials with
elastic–perfectly plastic behavior. It is assumed in the model that
the ultimate load-carrying capacity is reached when two plastic
hinges are formed, one in the concrete member and the other in
the timber member. The results obtained with this model were
compared with test results showing good agreement but making
an extrapolation to other conditions difficult due to the limited
amount of experimental results available. Besides, the consideration
of perfect plastic behavior for concrete may not always
be the best solution since concrete characteristically has brittle
behavior and some crushing will always occur under a steel fastener,
which was not accounted for in the model.
In this paper, the load-carrying capacities obtained from a
large number of laboratory tests with timber–concrete joints with
dowel-type fasteners are presented. The joints used various configurations,
including different materials concrete and timber,
and in some cases a floor board interlayer between the timber and
concrete. This is followed by the presentation of three models to
predict the load-carrying capacity of this type of joint. Each of the
models proposes a different approach to simulate damage on the
concrete side. The objective is not only to present test data for
dowel-type fasteners in timber–concrete joints but also to propose
valid models to predict load-carrying capacity with accuracy,
allowing for different timber species and concrete mixtures.
Experimental Program
Test Series Properties
The dowel-type fastener used on this work consisted of a dowel
produced from short pieces of steel reinforcement bars. Smooth
bars were used in six test series, and profiled bars in the rest. The
experimental program of short-term tests consisted of eight different
test series with various materials and configurations in
order to cover a large number of practical application cases.
Each one of these test series was thought to be used in certain
practical conditions always with a concrete slab over timber
beams deck systems with this joint type are not that common.
Dowel-type fasteners with normal concrete are commonly used in
practice, so it was decided to evaluate whether the profile of the
fastener has an influence in the mechanical behavior by testing
smooth and profiled fasteners. Lightweight aggregate concrete is
a good alternative in situations of renovation when the supporting
walls or the foundations do not have enough strength to support a
large extra load. On the other hand, high-strength concrete is an
option to decrease the thickness of the concrete slab by increasing
the mechanical performance of the slab, and at the same time the
mechanical properties of the joints. This may be important in
situations where there are dimensional limitations, for instance, in
slab thickness.
Chestnut and Maritime pine were selected because they are
traditionally used in timber floors in Portugal. Therefore, they can
be found in situations where timber–concrete systems are an interesting
option in the rehabilitation of these timber floors. In the
Steel
Table 1. Test Series Properties
Test
series Timber Concrete a f y N/mm 2 Fastener
8 mm Spruce C25/30 456 Smooth bar
10 mm A Spruce C25/30 496 Smooth bar
HSC Spruce C50/60 456 Smooth bar
MP Maritime pine C25/30 511 Smooth bar
C Chestnut C25/30 477 Smooth bar
LWAC Spruce LC16/18 462 Smooth bar
10 mm B Spruce C30/3 S500 b Profiled bar
INT Spruce C30/3 S500 b Profiled bar
a Concrete class in accordance with Eurocode 2 CEN 1997.
b Ultimate strength of the steel was assumed as the characteristic value,
obtained in accordance with Eurocode 2 CEN 1997.
same situation it is likely that boards from the timber floor are
used in the renovation process as a lost formwork, in that case,
the mechanical behavior of the joint will be different. This was
the reason for having one test series with an interlayer made of
floor boards. The properties of each one of the test series are
presented in Table 1.
Maritime pine and Chestnut specimens were produced from
solid timber, while Spruce specimens were produced from glued
laminated timber. Glued laminated timber was produced with
solid timber of Strength Class C18 according to CEN-EN 338
CEN 1995. Chestnut and Maritime pine were selected so that no
large knots or other defects were present in the area around the
fastener. The mean values of the density for these three wood
species was 454, 605, and 566 kg/m 3 , for Spruce, Maritime pine,
and Chestnut, respectively, with corresponding standard deviations
of 40.4, 92.6, and 51.9 kg/m 3 . The specimens were stored at
climate conditions 20/65 20°C temperature and 65% air humidity,
leading to a timber moisture content of around 12%. The
floor boards composing the interlayer were made of spruce planks
with 20 mm thickness and variable widths. Each one of the
boards was connected to the timber beams by two nails commonly
found in existing floors.
In a number of test series 8 mm, 10 mm A, C, LWAC, and
MP, when the shear test was finished, the timber was used to
produce embedment test specimens in accordance with CEN-EN
383 and to determine the material density. The material for these
tests was taken from an area next to the fastener. This procedure
was found to be the best way to determine the properties of the
timber used in the shear tests. In the other series, the embedding
properties were not measured while the density was measured in
representative samples.
The basic composition of all the concrete mixtures was similar:
sand, gravel, Portland 42.5/52.5 cement, water, and superplasticizer.
The differences between concrete classes were created by
the use of expanded clay instead of gravel in the lightweight
aggregate concrete and the addition of silica fume for highstrength
concrete. In practice, the thickness of the concrete slab is
small, generally between 30 and 100 mm. This not only increases
the difficulty of the vibration but also requires extra effort to
avoid undesired problems that can occur during the building
process, such as discontinuity in concrete introduced by large
aggregate size compared to the member thickness or difficulties
ensuring the dimensions of the member. In order to ensure better
quality of the joints, all the compositions used were prepared to
have fluid concrete and a maximal size of the aggregate of
12.7 mm. This corresponds to the recommended procedure at the
JOURNAL OF STRUCTURAL ENGINEERING © ASCE / MAY 2007 / 721

Fig. 1. Test specimen configuration and test setup dimensions
in mm
building site. The properties of concrete were evaluated from its
density and compression strength, estimated based on the weight
and compression strength of 150 mm cubes. This was first done
after 1 or 3 days, and then at 7, 14, 28 days and for the tests series
8 mm, 10 mm A, HSC, MP, C, and LWAC also on the day of
the short-term tests.
The dowels were produced from smooth bars of unknown steel
grade and profiled reinforcement steel bars from steel quality
S500 in accordance with Eurocode 2 CEN 1997. No treatment
to prevent steel corrosion was applied and some of the dowels
already showed superficial corrosion at the time of application.
Any dust or grease on the surface of the fasteners was removed
before they were driven into the timber. The strength of the steel
fasteners was evaluated in two different ways: when the steel
grade was known by the characteristic value given in the standards,
and when the steel grade was not known by means of one
tension test for each steel bar used.
Test Specimen Configuration
Prior to the test specimen geometry decision, two preliminary
tests series were performed Lopes et al. 2003; Dias et al. 2004a.
The main objective of the preliminary tests was to settle the adequate
test configuration, ensuring that the test results obtained
in the shear tests were representative of the phenomena developed
in the composite beams.
The test specimens consisted of one central concrete member
connected to two side timber members, as can be seen in Figs. 1a
and b for the specimens with floor boards. In the 10 mm B and
INT test series, the width of the concrete member was larger
420 mm, while for the rest the test configurations were similar.
Preparation of the timber elements for the shear specimens
began either with nailing the floor boards or, when there was no
interlayer, by predrilling the fastener holes. The predrilling was
done to the nominal diameter of the steel fastener and a depth
equal to the length of the fastener in the timber member. The
fasteners had depths inside the concrete of 60 mm 10 mm B and
INT and 40 mm the other test series, and inside timber of
80 mm 8 mm, 120 mm 10 mm B, INT and 100 mm the other
test series. This was followed by the application of either a plastic
foil or paint around the timber in order to avoid water takeup.
Furthermore, this procedure was also useful in reducing timber–
concrete friction during the tests, since the whole load is supposed
to be transferred by fastener shear. Finally, steel dowels were
hammered into the timber member. This corresponds to the recommended
procedure at the building site.
During the entire process, the timber specimens were stored at
a 20/65 climate 20±2°C temperature and 65±5% air relative
humidity. They were then placed in the formwork and cast. They
cured for three days in the formwork, followed by storage until
the day of the test.
The test setup was similar for all the test series, the load was
measured using a load cell, and the slip measured using LVDTs.
The load was measured at the point of application on the top of
the concrete member, while the relative displacement between
timber and concrete was measured at four locations at the center
of the test specimens see Fig. 1. In test specimens with an interlayer,
the relative displacement between timber and floor
boards was also measured at the same location using another four
LVDTs. The displacement transducers LVDT and plates necessary
to measure the displacements were assembled by means of
screwing on timber and gluing on concrete.
In order to distribute the forces applied in the top of the concrete
element and in the bottom of the timber members, thick
steel plates were used for the entire area. In Test Series 10 mm B
and INT, the bottom of the timber members were clamped together.
The other test series had no restrictions to the horizontal
movement, except the natural timber–steel friction. In the tests,
all the procedures given in CEN-EN 26891 CEN 1991 were
followed. According to this standard, the load-carrying capacity
of the joints is considered as the maximum load achieved up to a
slip of 15 mm.
Test Results and Analysis
Test Results
The test results together with some material properties from each
test series are presented in Table 2. As shown above, four
displacements were measured in each test. However, only the
average values are presented in Fig. 2. These four measurements
presented some differences, particularly between both faces of the
test specimen front and back due to the rotation of the concrete
member caused by small eccentricities.
The shear tests showed a large plastic deformation capacity for
all the series, which came in line with what was observed in the
preliminary tests Lopes et al. 2003; Dias et al. 2004a. In four of
the tests 10 mm B, three times; INT, one time, cracks showed
up in the concrete member in the plane of the fastener length. In
two of them this occurred before the 15 mm slip was reached
one in each of the series. In these two tests, the load was considered
as the maximum load achieved before the failure, which
occurred for slip values of 8.78 and 14.62 mm for Series
10 mm B and INT, respectively.
The behavior observed during these tests corresponds to the
yielding of the steel dowel in two points, leading to the formation
of two plastic hinges. On all occasions there were clear plastic
deformations in the steel fasteners, in the concrete immediately
next to the interface as well as inside the timber, indicating that
the two plastic hinges were at least partially formed. When the
timber was pulled from the concrete, it was possible to analyze
the damage in the concrete, which confirmed that there was also
damage on the concrete side around the dowel, mostly due to
crushing of the material. It also seemed that part of the dowel in
the concrete had no axial movement since the bond between the
722 / JOURNAL OF STRUCTURAL ENGINEERING © ASCE / MAY 2007

Table 2. Test Specimen Properties and Results
Test
series
Number
of tests
Number
of fasteners
per test
Concrete
Steel
f u N/mm 2 f c N/mm 2 f cc N/mm 2
Timber
f h N/mm 2
Age
days
Load per fastener
At 5 mm
kN
At 15 mm
kN
F max COV
15 mm
8 mm 21 2 456 40 160 41 193 5.9 6.8 0.05
10 mm A 21 2 496 46 184 44 190 9.6 11.3 0.10
HSC 21 2 456 84 336 44 187 9.7 11.8 0.07
MP 21 2 511 44 176 51 183 10.5 12.8 0.08
C 21 2 477 48 192 50 176 10.4 13.1 0.07
LWAC 21 2 462 27 108 39 55 7.8 9.3 0.04
10 mm B 10 4 S500 48 192 32 a 212 13.8 17.2 0.09
INT 10 4 S500 45 180 32 a 176 11.7 15.8 0.05
a Determined from the timber density using the formulation proposed in Eurocode 5 CEN 2003.
steel fastener and concrete was found to be perfect. Sometimes,
cracks appeared in the timber; however, they did not seem to
influence the load–slip behavior.
Corrosion on the steel dowels hardly increased from the time
they were placed; this is particularly true for the part inside the
concrete. The only exceptions were the dowels in the Chestnut
timber, which reacted with the timber and were much more corroded
than before manufacturing of the specimens. For that reason,
the pullout force necessary to remove the fastener from the
timber on that test series was much higher than the one necessary
to remove the fasteners of the other test series.
Analysis of the Results
At the time of the tests, the lightweight aggregate and normaland
high-strength concretes had compression strengths of around
27, 46, and 84 MPa, respectively. Since these concrete types
cover a wide range of compression strengths, any relevant influence
of concrete strengths on the behavior of the joints should
become clear. The values found for the load-carrying capacity
were 9.3, 11.3, and 11.8 kN per fastener for lightweight aggregate
and normal- and high-strength concretes, respectively see
Fig. 3. These results indicate that the load capacity increases with
the compression strength of the concrete. The greater difference
between lightweight aggregate concrete and normal concrete may
indicate that the influence is greater for concretes with lowerstrength
properties. The explanation for this may be in the fact
that for concretes with higher strengths the damage in concrete is
small, and thus becomes irrelevant when compared to the damage
in timber. In other words, the concrete acts as an almost perfect
constraint; therefore, the damage on timber has relatively higher
importance. Looking at the results, it can be concluded that the
load-carrying capacity of the test specimens with Maritime pine is
about equal to the ones with Chestnut with just 3% difference. On
the other hand, the load-carrying capacity of the specimens with
Spruce 10 mm A is almost 14% lower than the one from the test
specimens with Chestnut. These differences relate to the embedding
properties of the various timber species because no differences
in concrete behavior could be observed after the tests. The
higher values of the mechanical properties of Chestnut joints
compared to Maritime pine joints may be related to the higher
friction between timber and steel in the Chestnut series and to the
different characteristics of softwoods and hardwoods.
It is well known that the presence of an interlayer decreases
the load-carrying capacity Van der Linden 1999; Gelfi et al.
2002. On the shear tests presented here, the value of the loadcarrying
capacity decreased by around 8% when a floor board
interlayer of 20 mm was used.
Test Series 10 mm A and 10 mm B had similar characteristics
except for the dowel. In the first case, it was made of smooth steel
with an ultimate strength around 500 MPa, and in the second one,
of a profiled bar from the Steel Grade S500 assuming no influence
from the penetration depth once in both cases is higher than
10 times the diameter. Comparison between the load-carrying
capacity from both test series shows that the values of Test Series
10 mm A were only 66% of those in Tests Series 10 mm B. Furthermore,
the load–slip curves show that the load by which the
fastener yields is higher, but that the increase in the load after
yielding is also higher. The explanation of these differences may
be in the steel quality and surface shape of the fastener because a
higher steel strength results in a higher yield load. In addition, a
Fig. 2. Average load slip curves obtained in the tests
Fig. 3. Compression strength of concrete versus load carrying
capacity obtained
JOURNAL OF STRUCTURAL ENGINEERING © ASCE / MAY 2007 / 723

F e,c = 4My f h d
2
The last possibility, linear–elastic up to the failure with crushing,
is simulated assuming zero concrete strength in crushed areas.
The remaining part of the concrete foundation can be assumed
with either elastic–plastic or linear–elastic behavior. Here, the last
possibility was selected because the plastic deformation capacity
of concrete is small. Therefore, outside the crushed area the behavior
of concrete is probably more similar to linear–elastic than
to elastic–plastic. Eq. 3 gives the failure load for the joint assuming
the formation of two plastic hinges
F cr,c = df h− e +e 2 + 4M y
df h
3
Fig. 4. Material models assumed for concrete
profiled shape probably increases the contact forces between timber
and steel, resulting in a higher pullout resistance friction
along the fastener surface, increasing the shear forces transmitted
by the joint.
Models to Predict the Load-Carrying Capacity
The models presented here were obtained using a similar approach
to the one used to obtain the Johansen models Johansen
1949 for timber–timber joints, also known as the European yield
theory. The method assumes elastic–perfectly plastic behavior in
the timber and fastener. The load is calculated assuming a certain
failure mode; the final load is the lowest load obtained considering
the various failure modes. Joints with interlayer floorboards
have been presented in Dias et al. 2004b.
In order to apply the method to timber–concrete joints it is
necessary to define the behavior of concrete. The actual mechanical
behavior of concrete is characterized by an elastic stage
followed by a short phase where a certain amount of plasticity
occurs followed by material failure softening. This behavior
typically leads to some crushing of concrete under the steel fasteners
in timber–concrete joints. Three possibilities have been
chosen to simulate the nonlinear concrete behavior: elastic–
perfectly plastic, linear–elastic up to failure; and linear–elastic up
to the failure with crushing in concrete see Fig. 4.
The first possibility was analyzed by Gelfi et al. 2002, who
delivered a model to calculate the load-carrying capacity of joints
with and without an interlayer Eq. 1. Eq. 1 assumes that the
fastener is slender enough in relation to the embedding strength of
the materials to lead to the formation of two plastic hinges, one in
timber and one in concrete
F p,c = f h d 2
1+ 2M y
f h d + t 2
1+ 2 −
1+ t 1
If a linear–elastic behavior is considered, the equations are similar
to the equation found when analyzing steel–timber joints assuming
a thick steel plate and can be found, for example, in Eurocode
5 CEN 2003. In that case, the load-carrying capacity for a
failure with two plastic hinges depends only on the embedding
strength of the timber and on the fastener properties Eq. 2.
From the three proposals, in principle, the last one describes the
actual behavior with more accuracy. However, the crushed area of
concrete must be supplied as input to the problem, constituting a
significant limitation because usually that parameter is not known
in advance.
In this study, the load-carrying capacity was calculated per
fastener using the three models. The total load was then obtained
by the sum of the single loads for each one of the fasteners. The
models require the values of a number of material and geometrical
properties, namely, the embedding strength of the materials
and the yield/ultimate bending moment capacity and diameter of
the fasteners. Considering the data available from the tests, the
following assumptions were made in order to have the necessary
input values:
• The embedding strength of timber was obtained directly from
the embedment tests;
• The embedding strength of concrete was considered to be
equal to the compression strength of confined concrete estimated
according to indications from CEB-FIP MC90 CEB-
FIB 1990;
• In Eq. 3 the length of crushed concrete was assumed to be
equal to 5 mm based on analysis of tested specimens;
• The fastener diameter was assumed always to be equal to the
nominal diameter of a 10 or 8 mm steel bar;
• The ultimate bending moment capacity of the fastener was
calculated with Eq. 4 using the mean value of the ultimate
tension strength measured in the tension tests
M y = f u
600 180d2.6 4
Comparison between Experimental and Model
Results
The three models used here, neglect phenomena like the increased
friction between the members and the supports and geometric
nonlinearities, which are usually considered as limitations
Patton-Mallory et al. 1997. In order to estimate the influence of
these parameters, a 2D FEM model had been developed Dias
2005. From this model it was concluded that the ultimate loadcarrying
capacity, solely on the basis of embedding, steel, and
concrete strengths was reached between 3 and 6 mm displacement.
In that case, the values obtained with the equations proposed
better represent the loads obtained for 3–6 mm slip than
the loads obtained for 15 mm slip as stated in CEN-EN 26891
724 / JOURNAL OF STRUCTURAL ENGINEERING © ASCE / MAY 2007

Table 3. Ratios and Correlations between Model Results and Tests Results
Test series 10 mm A Chestnut HSC LWAC MP 8 mm Mean
a Correlation
F p,c 0.924 0.717 0.756 0.641 0.911 0.775 0.788
F ee 0.907 0.712 0.763 0.638 0.898 0.775 0.782
F cr,c 0.929 0.718 0.755 0.647 0.914 0.775 0.790
K ser EC5 0.326 — 0.656 0.457 0.702 0.444 0.497
b Ratios
F p,c 0.94 0.90 0.95 1.03 0.93 0.96 0.95
F ee 1.07 1.03 1.01 1.20 1.06 1.08 1.07
F cr,c 0.86 0.82 0.81 0.97 0.84 0.84 0.86
K ser EC5 1.20 0.79 1.29 1.21 1.11 1.19 1.13
Note: EC5Eurocode 5.
CEN 1991. For this reason, here, the model results are compared
with the load obtained for 5 mm slip, the values of which
are presented in Table 2.
In Table 3, the ratios between the results obtained with the
three models and the experimental results obtained from the tests
are presented, together with the correlation coefficients for these
two sets of data. The model and experimental results are plotted
in Figs. 5–7 for each one of the models proposed. The stiffness of
the joints was compared with the ones calculated in accordance
with the Eurocode 5 CEN 2003 model and were found to be in
reasonable agreement Table 3.
The best correlation between test results and model results was
obtained with the model assuming the behavior of the concrete as
linear–elastic with crushing. The exception is the test series with
high-strength concrete in which a higher correlation was obtained
with the model assuming linear–elastic behavior for concrete
F e,c . Except for the HSC test series, the worst correlations were
obtained with the model assuming linear–elastic behavior for
concrete.
Analysis of the ratios between the model and experimental
results shows that the model assuming elastic–perfectly plastic
material behavior for concrete predicts the test results best F p,c ,
but the difference from the one assuming linear–elastic material
behavior is small. The differences are on average 5% for F p,c ,7%
for F e,c , and 14% for F cr,c respectively. F p,c and F cr,c tend to
underestimate the load-carrying capacity, whereas F e,c tends to
overestimate it. The model with higher correlation coefficients is
the one with the highest differences between model results and
test results. Part of the explanation for this can be related to the
load transmitted by friction between timber and concrete and between
timber and steel. This effect is not considered in the models
but it has an influence on the load transmitted, especially when
the slip increases, as was shown by Dias 2005.
It is also interesting to verify that the best predictions for the
series with high-strength concrete were obtained using F e,c assuming
concrete as linear–elastic, and the worst for F cr,c assuming
linear–elastic with crushing. This indicates that when highstrength
concretes are used more than 84 N/mm 2 , in this case,
the concrete acts as almost a perfect clamp to the steel fastener.
On the other hand, when the ratios for lightweight aggregate concrete
are analyzed they show that closer predictions are obtained
with F p,c and the worst with F e,c . This indicates that when the
strength of concrete decreases, the assumption of perfect clamping
is too crude and the damages on concrete must be considered
in someway by assuming nonlinear material behavior.
The graphs confirm the tendency from the model F e,c to overestimate
the actual load-carrying capacity, and the models F p,c
and F cr,c to underestimate it, particularly the last one.
Example Calculation
As an example, an old timber floor is considered with Maritime
pine floor beams with a cross section of 80 mm160 mm spaced
by 350 mm. These structures need to be renovated in order to
increase the load-carrying capacity ultimate load 5 kN/m 2 and
service load 4 kN/m 2 . In addition, a timber–concrete composite
solution is assumed with a concrete layer of 40 mm thickness
made with C25/30 concrete in accordance with Eurocode 2
CEN 1997, connected to the timber beams by means of smooth
Fig. 5. Load per fastener obtained in the tests against the value
determined with Eq. 1 F p,c
Fig. 6. Load per fastener obtained in the tests against the value
determined with Eq. 2 F e,c
JOURNAL OF STRUCTURAL ENGINEERING © ASCE / MAY 2007 / 725

Fig. 7. Load per fastener obtained in the tests against the value
determined with Eq. 3 F cr,c
steel bars with 10 mm diameter spaced by 100 mm.
Using the parameters and results presented above and the indications
given in Eurocode 5 CEN 2003, the values of the
stiffness for the timber, concrete, and composite sections are
EI w =3.2810 11 N/mm 2 ; EI c =5.6910 10 N/mm 2 ; and
EI ef =1.2010 12 N/mm 2 . The correspondent design stresses in
timber and concrete are c,w =2.10 N/mm 2 ; t,w =10.47
N/mm 2 ; and c,c =7.82 N/mm 2 . t,c =0.17 N/mm 2 . The
deflection at midspan is ms =6.43 mm. The ratios between the
stresses and deformations that would occur in the original
structure only timber and in the composite timber–concrete
structure are ratio− c,w =11.0 −; ratio− t,w =2.2 −; and
ratio− ms =3.9 −. The composite structure has much lower timber
stresses than the original timber floor, particularly compression
stresses, which are carried mostly by the concrete in the
composite structure. The midspan deformation is significantly
lower as is clear from the ratio between them. The renovation of
this floor using a composite solution meets the design criteria that
were not met with the original timber floor.
Conclusions
Analysis of test results of joints with various concrete types
showed that the load capacity of joints increases with compression
strength. The increase is more significant for lower strength
concretes. Indeed, when high-strength concretes are used, the
concrete acts almost as a perfect clamp for the fastener and no
further increase in load-carrying capacity is observed. It was also
found that an increase in timber density leads to a significant
increase in the load capacity of joints. This supports the use of
traditional models for the load-carrying capacity of timber joints
where density is one of the main parameters. In terms of the
fastener used, it was found that the load capacity of joints increases
not only with the strength of steel but also with the shape
of the bar. Profiled bars showed higher increases in the load in the
yielding stage than smooth bars.
Comparison between model and test results showed that the
three models are able to predict the load capacity of joints with
reasonable accuracy. The best correlations between model and
test results were obtained with the model assuming linear–elastic
behavior with crushing for concrete, while the worst correlations
were obtained with the model assuming linear–elastic behavior
for concrete. The only exception was the test series with highstrength
concrete, for which the best correlations corresponded to
the model assuming linear–elastic behavior for concrete.
The most accurate predictions lower deviations between test
and model results were obtained for the model assuming elastic–
perfectly plastic behavior for concrete, which on average underestimate
the test results by 5%. The predictions of the model
assuming linear–elastic behavior for concrete were also close to
the test results; however, they overestimate them on average by
7%. The model assuming linear–elastic behavior with crushing
for concrete gave the worst predictions, in spite of having the best
correlations.
In general, this study gave indications that the model describing
better the phenomena involved takes crushing of concrete into
account. It leads to higher correlations between model prediction
and experimental results. However, the actual load-carrying capacity
seems better predicted by the elastic–plastic model, indicating
a larger standard “error” in the crushing model, deviating
in the prediction from the experiment. For joints where highstrength
concretes are used, the model with linear–elastic behavior
for concrete is a good and valuable simplification. However, it
significantly overestimates the load capacity of joints when lowerstrength
concretes are used.
Notation
The following symbols are used in this paper:
8mm test series with 8 mm smooth bar, glue
laminated Spruce, and medium strength
concrete;
10 mm A test series with 10 mm smooth bar, glue
laminated Spruce, and medium strength
concrete;
10 mm B test series with 10 mm profiled bar, glue
laminated Spruce, and medium strength
concrete;
C test series with 10 mm smooth bar, Chestnut
timber, and medium strength concrete;
d diameter of the fastener mm;
EI c stiffness of the concrete member N/mm 2 ;
EI ef stiffness of the composite member considering
a flexible joint system N/mm 2 ;
EI w stiffness of the timber member N/mm 2 ;
e length under the fastener with crushed
concrete mm;
F max load-carrying capacity obtained in the test in
accordance with CEN-EN 26891 kN;
F p,c load-carrying capacity for the model assuming
elastic perfectly plastic material behavior for
concrete kN;
F cr,c load-carrying capacity for the model assuming
linear–elastic behavior with crushing for the
concrete kN;
f c embedding strength of concrete N/mm 2 ;
f cc embedding strength of confined concrete
N/mm 2 ;
F e,c load-carrying capacity for the model assuming
linear–elastic behavior for concrete kN;
f h embedding strength of timber N/mm 2 ;
f u ultimate strength of the steel N/mm 2 ;
HSC test series with 10 mm smooth bar, glue
laminated Spruce, and high-strength concrete;
INT test series with 10 mm profiled bar, glue
laminated Spruce, medium-strength concrete,
and a 20 mm thick interlayer;
726 / JOURNAL OF STRUCTURAL ENGINEERING © ASCE / MAY 2007

K ser EC5 slip modulus calculated in accordance with
Eurocode 5 kN/mm;
LWAC test series with 10 mm smooth bar, glue
laminated Spruce, and lightweight aggregate
concrete;
MP test series with 10 mm smooth bar, Maritime
pine timber, and medium strength concrete;
M y plastic bending moment of the fastener
N/mm;
t thickness of the interlayer mm;
f c / f h ;
ms midspan deformation mm;
c,c design compression stress on concrete
N/mm 2 ;
c,w design compression stress on timber N/mm 2 ;
t,c design tension stress on concrete N/mm 2 ;
and
t,w design tension stress on timber N/mm 2 .
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