reforço de lajes fungiformes com recurso a …...with hydro demolition with high-pressure water...

Reforço de Lajes Fungiformes com Recurso a

Lâmina de Betão Complementar

Hugo D. P. Fernandes

RELATÓRIO 5

HiCon - Uso Racional de Betão de Elevada Resistência em Estruturas de Laje

Fungiforme Sujeitas a Ações Cíclicas e Sísmicas

(EXPL/EC M-EST/1371/2013)

Outubro de 2015

Reforço de lajes fungiformes com recurso a lâmina de betão complementar

II

INDICE

1. Introdução ...................................................................................................................... 1

2. Motivation ...................................................................................................................... 2

3. Analysis for unidirectional slabs strengthened with BCO .................................................. 2

Experimental programme....................................................................................................... 4

Definition of the strengthened specimens ............................................................................. 4

Test setup ........................................................................................................................... 6

Materials characterization .................................................................................................. 7

Experimental results .............................................................................................................. 8

Failure modes ..................................................................................................................... 8

Debonding and failure loads ...............................................................................................10

Discussion .............................................................................................................................12

Conclusions ...........................................................................................................................17

Acknowledgements ................................................................................................................18

References ............................................................................................................................19

4. Pulloff testing for characterizing the interface tensile resistance .......................................21

Motivation and testing ...........................................................................................................25

Results and discussion ...........................................................................................................30

Conclusions ...........................................................................................................................33

References ............................................................................................................................35

5. Numerical analysis for unidirectional flat slabs strengthened with BCO ...........................37

Introduction ..........................................................................................................................37

Concrete-to-concrete interface ...............................................................................................38

3D Specimen modelling .........................................................................................................40

Results ..................................................................................................................................44

Discussion .............................................................................................................................49

Conclusions ...........................................................................................................................53

6. Punching tests for bi-directional reinforced flat slabs strengthened with BCO ..................54

Experimental research ..........................................................................................................54

Strengthened test specimens definition ..........................................................................54

Test setup ........................................................................................................................57

Materials characterization ..............................................................................................59

Experimental results .............................................................................................................60

Failure modes ..................................................................................................................60

HiCon - Uso Racional de BER em Estruturas de Laje Fungiforme Sujeitas a Ações Cíclicas e Sísmicas

III

Debonding and failure loads............................................................................................62

7. Thesis and Dissertations .................................................................................................65

Università Degli Studi Firenze ...............................................................................................65

8. Communications in conferences ......................................................................................66

SILE2015 ..............................................................................................................................66

fib – Cape Town 2016 ............................................................................................................67

9. References .....................................................................................................................69


1

1. Introdução

O presente relatório vem enunciar e descrever as tarefas realizadas pelo aluno de Doutoramento Hugo

Daniel Pereira Fernandes, durante o ano letivo de 2014/2015. O mesmo consistiu na análise de lajes

fungiformes armadas numa direção e produção de um artigo científico sobre o trabalho experimental

resultante das mesmas, realização e análise de resultados decorrentes dos ensaios de lajes armadas em

duas direções, concebidas no período anterior desta bolsa, e definição/conceção de novas lajes da mesma

geometria. Foi também realizada análise numérica através de modelação numérica não linear das lajes

retangulares e quadradas. Os trabalhos de dimensionamento e conceção de modelos de laje fungiforme

consistiram no desenvolvimento de quatro modelos de laje fungiforme, armados em duas direções,

baseando-se nas conclusões e ilações de relatórios anteriores.

O programa experimental executado no decorrer deste terceiro período consistiu em ensaios de

punçoamento monotónico centrado de quatro modelos de laje armados em duas direções, e ensaios do

tipo Pull-off em provetes com 150 mm de aresta. Os mesmos permitiram aferir o modo de rotura da

interface, sendo esta altamente solicitada por esforços de corte e de tração decorrentes dos fenómenos de

punçoamento, bem como para a peça inteira, sendo possíveis a rotura por flexão ou punçoamento.

A instrumentação utilizada visou a monitorização dos esforços desenvolvidos no decorrer dos ensaios, a

deformação da laje reforçada, tendo sido ainda monitorizados os assentamentos dos apoios, a deformação

da camada de reforço em relação ao substrato, bem como medidas as extensões das armaduras das duas

camadas, com vista à análise da diferença das extensões entre ambas, e estimativa dos esforços

desenvolvidos na interface.

Foram realizados quatro ensaios de punçoamento monotónico de modelos de laje reforçados com uma

nova camada de betão na zona tracionada, divididos em modelo de referência e três soluções de acordo

com relatórios anteriores: com conectores de corte distribuídos na interface, amarração das armaduras da

camada de reforço nas extremidades, e outro onde se combinaram as anteriores. Foram ainda

contabilizados cento e dois ensaios do tipo Pull-off, em provetes moldados in-situ nas extremidades das

lajes.

Durante a realização destes ensaios, houve a contribuição de três alunos da Universitá Degli Studi

Firenze, Massimo Lapi, Daniele Martini, e Emilio Zagli, durante todo o programa experimental. Estes

alunos contribuíram positivamente na realização dos ensaios experimentais, tendo ainda contribuído com

uma análise de resultados que fruiu na obtenção dos respetivos graus de Mestres em Engenharia Civil

pela mesma instituição.

Devido ao estado avançado dos trabalhos das duas cadeiras constantes do plano de doutoramento,

“Comportamento de Materiais Estruturais” – sob a orientação do professor Carlos Chastre, e “Elementos

Finitos em Engenharia de Estruturas” – sob a orientação do professor Cornélio Cismasiu, pretende-se que

terminem durante o segundo semestre do ano letivo 2015/2016. Também a cadeira de “Estudos

Avançados de Betão Armado” – sob a orientação do professor Válter Lúcio, terminará durante o mesmo

semestre.

Devido à natureza da abrangência pretendida com a comunicação deste trabalho, o presente relatório será

doravante redigido em língua Inglesa.


2

2. Motivation

The work produced during this twelve-month period follows the conclusions from previous reports

(REF), where twelve unidirectional reinforced concrete slabs were tested. These slabs allowed for the

definition and detailing of new bi-directional reinforced flat slabs, which were tested in a punching test

setup, standard in NOVA University of Lisbon, with a few improvements for better understand the

phenomenon between layers.

Since we’re dealing with a tension phenomenon, pull-off tests were defined to attest the interface tensile

resistance. The results from these tests would allow for correct definition of numerical parameters needed

for correct prediction of the strengthened specimens behaviour.

From previous tests we’ve learned that the debonding phenomenon can rule the global behaviour of a

strengthened structure, bridging the local and global failure behaviour expected from the resulting

composite structure. Two-step behaviour, before and after debonding of the two layers, could be

identified from those tests and is expected in the new bi-directional tests.

The numerical modelling of the unidirectional slabs is assumed to identify the phenomena that occurs

mainly in the interface of the two layers, identifying important factors that govern the behaviour of the

global structure. These parameters are currently unable to be identified in experimental testing, with

numerical testing filling the gaps of knowledge of this specific behaviour.

The pull-off tests designed alongside the experimental tests in this work allowed for the definition of the

interface tensile resistance, and tackled some issues identified in previous works, mainly the verticality of

the applied force and the contact surface area which can govern the relationship between the concrete

larger aggregates and the specimen total area.

The aforementioned motivation will allow for the identification of the behaviour governing phenomena,

and determination of new punching tests, that will allow varying the geometry of the strengthened

specimens and correctly define the characteristics for structures strengthened with this solution.

3. Analysis for unidirectional slabs strengthened with BCO

Strengthening of concrete structures by adding a new concrete layer is well known when applied to the

compressed face of concrete elements. Examples of that are beams and columns strengthened with

concrete jacketing. When applied on the tensile face of concrete elements, bonding between two concrete

layers is more complex, since relative deformation will cause debonding of the added layer.

This technique relies on the quality of bond between the two concrete layers, therefore varying with

surface preparation, and if steel connectors are installed crossing the interface. If no connectors are used,

adhesion is the only component of the resisting mechanism acting on the interface, and brittle failure shall

occur. This relies strongly on roughness, which allows for interlocking of the two layers and consequently

bonding stresses to develop along the interface. With steel connectors crossing the interface between the

two concrete layers, three components of the resisting mechanism illustrated in Fig. 1 shall develop [1]:

1. Adhesion – due to chemical and physical bond between the two layers, and mechanical

interlocking shall be considered if macroscopic surface roughness is present;

2. Friction – direct consequence of external loads perpendicular to the interface, or due to steel

connectors crossing the interface that are mobilized in tension for higher relative displacements,

with equilibrium guaranteed by compressive forces at the interface;

3. Dowel action – which refers to the bending resistance of steel connectors crossing the interface

due to relative slip, though the tensile forces cause them to be simultaneously loaded in shear, and

thus not reaching flexural resistance of the steel connector.


3

Fig. 1 Resisting mechanism for concrete-to-concrete interfaces with steel connectors.

Bond performance of the interface will depend on the appropriate conditions for these resisting

mechanisms. Some authors provide recommendations for surface preparation and curing conditions of the

added concrete layer. Contamination of the concrete surface before casting the new layer, methods used

for surface preparation, and surface microcracking, are according to [2] the main quality parameters for

the application of this technique. A special attention to edge zones is also referred to in the document

since the discontinuity of the cross section allows for the development of significant tensile and shear

stresses.

Interface bond performance varies from null to full transfer of horizontal stresses between the two

layers, allowing monolithic behaviour to be achieved, according to Fig. 2.

Fig. 2 Interface performance regarding shear stress transfer between layers.

Adhesion/Interlocking + Friction Dowel Action

τ

τ σ

σ

slip


4

Usually interaction between layers in the cross section is characterized as partial, and relative

displacement between the two layers is accounted for. Several behaviour models have been developed

over the years ([3], [4], [5]), where relative displacement, both horizontal and vertical, is comprised.

Stress transfer between the two concrete layers is based on the shear friction theory ([6], [7], [8], [9],

[10]). The behaviour models developed using the basis of this theory account for relative slip and

interface crack opening, which are responsible for crushing of concrete and reduction of the contact area,

respectively.

Models for relative displacement analysis and some models for stress analysis do not specifically

account for steel connectors crossing the interface. This is a behaviour changing aspect for detailing the

interface since it severely limits interface crack opening, providing greater stresses to develop and larger

slips. Steel connectors crossing the interface, properly anchored to both layers to improve strength, can

reduce uncertainty about interface performance and should be accounted for in the design of these

interfaces [11].

Surface preparation also plays a major role on interface behaviour due to the interlocking mechanism,

which causes the interface crack to open when relative slip occurs [12]. There are several methods for

surface preparation, more or less intrusive in terms of microcracking of the existing concrete layer, which

can affect interface performance ([13], [14], [15], [16]). This can cause premature spalling of concrete

chunks and reduced interface performance, reaching a depth of 3 mm [17] to 10 mm [18] in the existing

concrete layer. Since a new concrete layer is to be cast against the existing one, a greater roughness is

required, with exposed aggregate particles for improved interlocking. This is quantified in terms of an

average roughness parameter, which can be assessed through several techniques. This value defines the

interface in terms of roughness, according to [19], from smooth to very rough, with a specific profile

geometry. For improved performance of the interface, a very rough profile is recommended, with an

average roughness parameter equal or greater than 3,0 mm, since it directly improves all other

components of the resisting mechanism.

Preparation of the surface prior to casting the new concrete layer is done using several methods ([13],

[14]). Chipping with an electric or pneumatic hammer and steel moil point is the most common, along

with hydro demolition with high-pressure water jet. The latter technique is the most efficient in terms of

roughening and reduced microcracking of the surface, but also characterized by difficult logistics and

higher costs. For this reason, chipping with an electric hammer and steel moil point is the easiest and least

specialized method for surface roughening in a strengthening/retrofitting situation. Values for adhesive

tensile stress are provided in [14] for the steel moil point and the high-pressure water jet of 1,10 MPa and

1,46 MPa, respectively.

Besides geometrical characteristics of the interface, performance of the strengthened structure also

depends on materials of both substrate and the new layer to be overlaid. They directly influence both local

and global behaviour, through adhesive capacity and crushing resistance of concrete, and through bending

and shear resistance of the composite layers. According to [20], the newly added layer should be of

greater resistance than the substrate and low shrinkage, and should be fluid enough to penetrate the

grooves on the existing surface.

Experimental programme Definition of the strengthened specimens

For characterizing the behavioural aspects of concrete layers in a strengthening situation, three different

detailings of the interface were tested. For reference, some specimens were tested with only surface

roughening. The experimental tests comprised then:

1. Surface roughening only (S-REF);

2. Steel connectors distributed along the interface, with 50 mm of anchorage length (S-STC, 1 to

3), and 70 mm of anchorage length (S-STC-4);


5

3. Anchoring of the longitudinal reinforcement bars of the new layer at the ends, since the

deformation of the specimens in bending suggests the lifting of the overlaid concrete in this zone

[21], with 50 mm of anchorage length (S-ANC, 1 to 3) and 70 mm of anchorage length (S-ANC-

4);

4. All of the aforementioned techniques combined for improving bond at the interface (S-STANC).

Also in the MC 2010 [19] the lifting phenomenon at the overlaid concrete edge is considered relevant

since deformation due to shrinkage is greatest in this zone and debonding shall occur. Roughening of the

substrate surface was accomplished for all specimens with an electric hammer and steel moil point.

The substrates were comprised of a rectangular cross section, with 1.00 m in width and 0.12 m in

height, reinforced with double 10 mm diameter longitudinal bars (Fig. 3). Reinforcement for the overlaid

concrete layer consisted of six double 12 mm diameter bars.

Fig. 3 Rebar detailing and strain gauge placing for reference specimens (S-REF).

The contact area between the two layers was reduced transversally to increase bond stress. This layer also

did not reach the supports for all specimens, since confinement of the rebar does not necessarily happen

in a real strengthening situation.

The overlaid concrete rebars were instrumented with strain gauges, as well as the substrate rebars at

midspan and 0.25 m from this point. Interface detailing with longitudinal rebars crossing the interface or

steel connectors with anchorage lengths of 50 mm and 70 mm can be observed in Fig. 4 through Fig. 6.

Fig. 4 Rebar detailing and strain gauge placing for S-ANC specimens.

Fig. 5 Rebar detailing and strain gauge placing for S-STC specimens.


6

Fig. 6 Rebar detailing and strain gauge placing for S-STANC specimens.

Surface preparation for casting the new layer was performed with an electric chipping hammer and steel

moil point. The final surface can be observed in Fig. 7. Despite the disadvantages concerning

microcracking of the existing surface, this method was chosen due to its practical and economical

characteristics. Surface roughness was evaluated for each specimen, with concern that the moil tip should

not go deeper than 10 mm. This protects the longitudinal rebar and the minimum surrounding concrete

cover.

Fig. 7 Substrate surface roughness assessment setup and detail.

Several methods for surface roughness assessment can be used [35], with the sand patch method being the

most widely referred in the literature. Due to the availability of other methods, and with knowledge from

[22] that this method is limited for very rough surfaces, another method by point measurement was used

to evaluate the roughness profile. An average surface roughness of 3.1 mm was determined for all

specimens. This value allows for a surface characterization according to MC 2010 [19] as very rough.

Test setup

The specimens were subjected to monotonic loading in a three point bending test, as shown in Fig. 8.

Loading was imposed at midspan with hydraulic jacks, and reactions at two symmetrical supports,

according to the figures below. Forces at the supports were measured with four TML CNC-200KNA load

cells and the deformation at the coordinates identified in the left picture below, with TML CDP-100

displacement transducers. These tests were performed with prestressing strands at the supports.


7

Fig. 8 General test setup.

For assessing relative displacements between the two layers, TML CDP-50 displacement transducers

were fixed on the substrate layer, as shown in Fig. 9.

Fig. 9 Relative displacement measurement setup.

For data acquisition, four HBM Spider8 datalogger units were used along with one HBM UPM100

datalogger unit, all monitored by HBM Catman V6.0 software. Loading was controlled by force with a

WALTER+BAI PKNS19D electronically controlled hydraulic pump, at a speed of 0.10kN/s for all tests.

Materials characterization

Material characterization performed for the different concrete layers, different rebar sizes, and grout used

for anchoring the rebar crossing the interface. The mechanical properties for the grout were tested

according to RILEM’s PC-5 [23] and PCM-8 [24] for the compressive and tensile strength, and with pull-

out tests for the bond strength. The parameters for this material were a flexural tensile strength of 9.7

MPa, a compressive strength of 78.8 MPa, and maximum bond stress of 16.2 MPa. The results for the

other materials can be observed below in Table 1. and Table 2. .

Table 1. Concrete strength characteristics.

Reference

(S-REF, 1 to 3)

Steel connectors

(S-STC, 1 to 3)

Anchorage

(S-ANC, 1 to 3)

Steel

connectors

(S-STC, 4)

Anchorage

(S-ANC, 4)

Steel connectors +

Anchorage

(S-STANC)


8

fc,cube

[MPa]

Substrate 45.3 40.6 37.4 55.5 54.8 56.4

Overlay 47.8 47.8 47.8 41.1 39.4 41.5

fct [MPa]

Substrate 2.8 2.5 2.3 3.9 3.9 4.0

Overlay 2.9 2.9 2.9 3.1 3.0 3.1

fc,cube – mean value for the compressive strength of concrete in cubic specimens.

fct – mean value for the tensile strength of concrete.

Table 2. Strength characteristics for the steel bars used.

Diameter Ø 6 Ø 10 Ø 12

fy [MPa] 541.0 530.6 532.1

fu [MPa] 692.7 627.5 627.4

fy – mean yield stress of steel.

fu – mean tensile strength of steel.

Experimental results Failure modes

Behaviour was consistent for all specimens in terms of cracking, debonding of the overlaid concrete, and

failure load. Cracking at the interface level for the reference specimens started from the ends of the

overlaid concrete layer and evolved along the interface to midspan, according to Fig. 10. Debonding

occurred when not enough contact area was guaranteed between the two layers, resulting in a stiffness

loss. Flexural failure then occurred for the substrate layer (Fig. 10, right).

Fig. 10 Interface cracking for reference specimens.

For specimens with longitudinal rebar anchored on the substrate layer, beginning of interface cracking

was similar to the reference specimens, occurring again at the overlaid concrete ends. Debonding

occurred very close to the failure load for these specimens. Flexural failure mode was not exclusive in

this case, with one specimen failing in shear (specimen S-ANC-1), as shown in Fig. 11.


9

Fig. 11 Failure for specimens with longitudinal rebar anchored.

Some specimens with smaller anchorage length (50 mm in average), had failure controlled by pullout of

the rebars anchored in the substrate, as shown in Fig. 12.

Fig. 12 Failure for specimens with small anchorage length of the longitudinal rebars.

For specimens with steel connectors distributed along the interface, cracking was similar to anchored

rebar ones, with two specimens failing in shear (see Fig. 13). There was also no visible full debonding of

the overlaid concrete, with failure occurring for the two layers resisting the load.

Fig. 13 Failure for specimens with steel connectors in shear (left) and bending (right).

Anchorage failure occurred for some specimens with pullout at the steel-grout interface visible through

the interface crack right after testing (Fig. 14, left). Through a longitudinal sectioning of the specimens,

the evolution of the shear crack for the specimen that failed in bending can be observed. The insufficient

anchorage length caused the shear crack to fail intercepting the steel connectors (see Fig. 14).

Shear failure

Interface crack


10

Fig. 14 Steel connectors anchorage detail and section for the S-STC specimens.

All three specimens with a larger anchorage length of rebars crossing the interface (70 mm) failed in

shear, as visible in Fig. 15. This attests the increase in performance when reinforcement crosses the

interface properly anchored, achieving an almost monolithic behaviour until failure.

Fig. 15 Shear failure for specimens with larger anchorage length of rebars crossing the interface (left,

S-ANC-4; centre, S-STC-4; right, S-STANC).

Visible cracking of the overlaid concrete focused near the midspan, according to Fig. 16 for the

specimens with rebar crossing the interface (S-ANC and S-STC). Reference specimens showed only

hairline cracks at midspan, which closed after unloading of the overlaid concrete layer after debonding.

Fig. 16 Cracking for the overlaid concrete at midspan (left, S-ANC, right, S-STC).

Debonding and failure loads

The relationship between load and deflection at midspan was analysed to characterize the behaviour of

the strengthened specimens. Figures 17 and 18 present the relationship between vertical load and

deflection at midspan, comparing specimens with reinforcement crossing the interface and reference ones.

The marker on the load-deflection curves illustrates the moment when the highest strain was registered

for the overlaid concrete rebars.


11

Fig. 17 Load-deflection curves for specimens w/ reinforcement crossing the interface anchored 50 mm. highest strain registered for the overlay concrete rebars

The load-deflection curves above show an approximately linear behaviour for all tests up to a load of 40

kN at midspan, followed by a reduction in stiffness. This was consistent with the cracking load for the

geometry of the cross section. For the reference specimens, debonding of the overlaid concrete occurred

for a load of 80 kN at midspan. Stiffness was reduced to approximately zero, observed graphically by an

horizontal plateau. This was followed by reloading of the substrate layer, until flexural failure occurs for a

load of 160 kN at midspan.

For the specimens with longitudinal rebar anchored 50 mm in the substrate (Figure 17, left), the failure

load was about the same as for the reference specimens. The debonding load of 142 kN represents an

increase of 79 % when compared to reference specimens. For specimens with steel connectors anchored

50 mm in the substrate (Figure 17, right), debonding load increased about the same as the latter (76 %),

with the resulting failure load 10 % higher than the reference tests. Debonding phenomenon was not

visible during testing, and only identified through the strain measurement at the longitudinal rebars.

The three specimens with 70 mm anchorage length differed from the first set of tests mainly due to

greater debonding and ultimate loads. The relationships between vertical load and deflection at midspan

for these specimens are presented below.

0

40

80

120

160

200

240

0 5 10 15 20 25 30 35

Load

at

mid

span

[k

N]

Displacement at midspan [mm]

S-ANC-1 S-ANC-2 S-ANC-3

S-REF-1 S-REF-2 S-REF-3

0

40

80

120

160

200

240

0 5 10 15 20 25 30 35

Load

at

mid

span

[k

N]


S-STC-1 S-STC-2 S-STC-3


0

40

80

120

160

200

240

0 5 10 15 20 25 30 35

Load

at

mid

span

[k

N]

Deflection at midspan [mm]

S-ANC-4 S-STC-4 S-STANC



12

Fig. 18 Load-deflection curves for specimens w/ reinforcement crossing the interface anchored 70 mm.

highest strain registered for the overlay concrete rebars

From the load-deflection curves, one can observe that the increase in anchorage length also increased

debonding and failure loads. The latter showed an increase of 39 % for the anchored longitudinal rebar

specimens, 27 % for the slabs with steel connectors, and 43% for all techniques combined. Also

noticeable is the small increase for the failure load of the specimen that combined all techniques (S-

STANC), when compared to anchored longitudinal rebar tests (S-ANC). This attests the impact of the

latter in the overall behaviour of the interface, due to longitudinal rebars larger diameter and anchorage

positioning at the overlaid concrete ends. The limitation in the evolution for both interface crack opening

and relative slip results in smaller relative displacements that reach the steel connectors, limiting its

contribution to the resisting strength of the interface. However, these are activated later in the loading, but

for smaller stresses.

The table below lists the failure loads for all strengthened specimens, and the load for which

maximum strain was reached in the overlaid concrete rebars, which is also the load for maximum stress at

the interface.

Table 3. Maximum strain at the overlaid concrete rebars and failure load at midspan.

Specimen

S-REF S-STC S-ANC S-

STANC 1 2 3 1 2 3 4 1 2 3 4

Max

imu

m s

trai

n

at

the

ov

erla

y r

ebar

s

Midspan load

[kN] 82.4 --- 80.1 140.3 --- 144.3 200.3 --- 150.0 140.0 190.3 214.7

Overlay rebars

Fso [kN] 138.9 --- 141.3 218.7 --- 226.5 375.9 --- 305.5 273.6 408.2 483.4

Substrate rebars

Fs [kN] 90.8 --- 88.6 179.7 --- 175.9 408.1 ---

153.6

137.3 272.7 235.2

Failure load at midspan

[kN] 163.1 155.3 160.8 177.2 171.2 180.1 201.8 151.1 153.6 157.3 221.2 227.7

Discussion The provisions on the Model Code 2010 [19] for concrete-to-concrete interface resistance were

considered for behaviour characterization, and analysing the components of the resisting mechanism,

aggregate interlock, friction, and dowel action.

Shear stresses at the interface are a consequence of the variation of force Fso along the longitudinal

rebar of the overlaid concrete, that are transferred to the substrate layer. These stresses are responsible for

the integrity of the composite section, since failure of the interface causes failure of the strengthened

element. Shear stress can be evaluated in terms of the force transferred between the two concrete layers

over its interface. In Fig. 19, the shear stresses acting on the composite section are shown for a finite

length l.


13

Fig. 19 Stresses for composite cross sections.

In a concrete element of length Δl between cracked sections, equilibrium of forces leads to an average

shear stress νi at the interface level. This is proportional to the variation of forces in length of the overlaid

concrete rebars, according to Eq. (1), where b is the width of the interface.

bl

Fv so

i.

(1)

The strain gauges installed in the longitudinal rebars provided a linear distribution of steel strains,

which is consistent with a uniform distribution of shear stresses across the interface. The values for the

steel strains at midspan εs and corresponding values for shear stress at the interface (νi) can be observed in

Table 4. These values are the maximum strains registered for the longitudinal rebars at the overlaid

concrete. Only two specimens of each detailing were instrumented with strain gauges at the rebars.

Table 4. Overlay steel strains and shear stress at the interface.

Specimen

S-REF S-STC S-ANC S-STANC

1 3 1 3 4 2 3 4 -

s [x10-6] 511.7 520.6 805.7 834.5 1384.9 1125.5 1008.0 1503.9 1780.9

νi [MPa] 0.46 0.47 0.73 0.76 1.25 1.02 0.91 1.36 1.61

Evaluation of the shear resistance at the interface was carried out according to Randl [2] and the MC 2010

[19]. The resisting mechanisms at the interface are divided in three main components: aggregate

interlock, friction, and dowel action of the reinforcement crossing the interface.

The low stiffness of slabs results in large vertical deflections and large horizontal displacements. In [2]

a distinction is made between stiff and more brittle behaviour of the interface in terms of the relative slip

s. This limit is set at 0.05 mm, governed by aggregate interlock and friction from external actions. For

slips over this limit, the behaviour is considered more ductile, with adhesion replaced by

friction/interlocking and dowel action. The following equations quantify both these scenarios, without

accounting for external actions that are favourable to the resisting mechanism:

σc+

Δσc

σc

σso σso+ Δσso

σs σs+

Δσs

M+ΔM M

Δl

νi

vi

ν


14

cdctdaadR ffcvmms 5,0,05.0 , (2)

cdccdydiydickrdowfrilR fffkfkfcvmms 21

3/1

,,05.0 (3)

Where:

vR,ad is the design value for the adhesive shear stress for the interface;

ra cc , are the coefficients for surface condition;

ctdf is the design value for concrete tensile strength;

is the reduction coefficient for compressive forces ( 3/1/30.55.0 ckf ≤ 0.55);

cdf is the design value for concrete compressive strength;

vR,il+fr+dow is the design value for shear stress of the interface for interlocking, friction, and

dowel action;

ckf is the characteristic compressive strength for concrete;

is a friction coefficient;

i is the ratio of reinforcement crossing the interface ( iisi AA /, );

k1 is the latter reinforcement performance reduction factor (yis fk /,1 ≤ 1.0)

where s,i is the actual tensile stress in the steel crossing the concrete interface;

ydf is the design value for steel yielding stress;

k2 is the interaction coefficient for flexural resistance of the rebar (≤ 1,6 for circular

cross-sections and C20/25 - C50/60);

c is the coefficient that accounts for the strut inclination of concrete in

compression.

An adjustment is proposed for some coefficients in an analysis situation that fits better with test

results. Coefficient cr is proposed as 0.2 for very rough surfaces. The shear strength due to aggregate

interlock for the reference specimens resulted in 0.66 MPa, calculated considering the average concrete

compressive strength. This value overestimates the test results of about 0.47 MPa for the two reference

slabs in Table 4. The tensile stresses that result from equilibrium at the end of the interface (Fig. 20) can

justify this behaviour, debonding prematurely for a brittle resisting mechanism.


15

Fig. 20 Vertical tensile stresses at the end of the interface.

The coefficient of friction μ was calculated as proposed by the authors in [25]. Considering the

average roughness presented in section 0, the friction coefficient takes the value of 1.4. For the coefficient

k1 a new method is proposed, that contemplates the amount of horizontal force at the interface not resisted

by dowel action of the rebar. This shear force at the interface (k-F ·Fso) is accounted for in the

quantification of the steel stresses for the rebars crossing the interface, by the means of a coefficient ‘k-F’

as follows:

si

iF

si

soF

si

cs

si

k

A

Fk

A

F

tantan,

(4)

Where Fs,c is the tensile force in the steel connectors or longitudinal rebar anchorage and ρsi is the

reinforcement ratio crossing the interface (Asi/Ai). θ is the angle between the interface plane and the

concrete strut that results from nodal equilibrium, as illustrated in Fig. 21.

Fig. 21 Resisting mechanism of the overlay rebars anchorage (left) and shear connectors (right).

The values for the coefficient k-F were determined according to the amount of stresses resisted by

dowel action of the rebar crossing the interface. Values around 70 % of the total horizontal load at the

interface for the steel connectors were determined for the aggregate interlock and friction resisting

mechanisms. The remaining stresses were then resisted by dowel action of the rebars. For the specimens

with longitudinal rebar anchored, the latter value was reduced by a factor of 0.5, since the edge lifting

phenomenon is present, thus resulting in tension of the anchored rebars. An angle θ of 21.8 º was

considered, with good correlation to test results.

Dowel action resistance alone can be calculated with Eq. (5), where maximum allowable dowel action

of the reinforcement is scaled down due to the interaction between bending and tensile stresses in the

rebars (√𝟏 − 𝒌𝟏𝟐), and due to opening (s) of the interface crack (√𝒔 𝒔𝒎𝒂𝒙⁄ ≤ 𝟏. 𝟎). This phenomenon is

particularly important for surfaces with a higher roughness, where horizontal relative displacement leads

to vertical displacement, causing the interface crack to open.


16

3

.1./.)( 2

1,

ysi

máxmáxFF

fAkssVsV (5)

Where:

ycubecsimáxF ffAkV ... ,2, is the maximum allowable force for dowel action;

siA is the area of the reinforcement crossing the interface;

cubecf , is the concrete compressive strength in cubic specimens;

yf is the steel yield stress;

s is the relative slip of the layers at the interface;

maxs is the relative slip for max,FV , limited to 0.10Ø - 0.20Ø.

The coefficient k2 is taken as 1.5, fitting within the values prescribed in [2], without safety factors. The

values for the resisting stress that results from dowel action, considering smax = 0.10Ø, are presented in

Table 5.

Table 5. Dowel action resistance.

Specimen

S-STC S-ANC S-STANC

1 3 4 2 3 4 (Ø6) (Ø12)

Asi [mm2] 336 678 336 678

s [mm] 0.48 0.53 2.38 0.51 0.49 1.32 1.69

k1 [-] 0.34 0.35 0.59 0.24 0.21 0.32 0.25

VF [kN] 62.5 62.2 60.6 91.1 91.7 140.1 72.7 146,7

νF [MPa] 0.21 0.21 0.20 0.30 0.31 0.47 0.73

νF/νi [-] 0.29 0.27 0.16 0.30 0.34 0.34 0.45

Contribution of the rebar crossing the interface is significant, accounting in the S-STANC solution for

almost one-half of the horizontal load at the interface by dowel action of the rebar. For most of the other

specimens, this mechanism accounted for around one third of the horizontal load at the interface.

The remaining resisting mechanisms at the interface can then be estimated with Eq. (6), and are presented

on Table 6. along with the total value for the resisting strength of the interface.


17

yicmfril fkfcv 1

3/1

r (6)

Table 6. Shear stress for friction and interlocking of protruding aggregates.

Specimen

S-STC S-ANC

S-STANC

1 3 4 2 3 4

νil [MPa] 0.64 0.64 0.62 0.63 0.64

νfr [MPa] 0.29 0.30 0.49 0.20 0.18 0.27 0.31

νfr+il/νi [-] 1.27 1.24 0.90 0.81 0.88 0.66 0.59

νfr+il+F [MPa] 1.13 1.14 1.33 1.12 1.11 1.37 1.69

νfr+il+F/νi [-] 1.55 1.51 1.06 1.10 1.21 1.00 1.05

An overestimation for the shear stress at the interface seems characteristic for the specimens with smaller

anchoring of the reinforcement crossing the interface. A good correlation with the model is found for the

specimens with proper anchoring of the reinforcement crossing the interface (S-STC-4, S-ANC-4, and S-

STANC).

Conclusions Twelve flexural tests were performed on slab specimens strengthened with a new concrete overlay.

Failure mode for each detailing of the interface was identified, along with several constraints to the

application of this strengthening technique. Full brittle debonding of the new layer occurred for the

reference specimens, attesting the importance of reinforcement crossing the interface. The smaller

dispersion of results attests the constant behaviour of concrete elements strengthened with this technique,

thus crediting its application in actual strengthening and retrofitting situations. Zones where the interface

surface could be observed were identified with adhesive or cohesive failure. The latter could be observed

generally in the substrate layer, where concrete was weakest.

One constraint identified after testing was the insufficient anchorage length of the reinforcement

crossing the interface in a first series of tests, which lead to a complementary experimental campaign.

Insufficient bonding length in the existing layer penalized the resisting strength of the interface, leading to

lower resistance and premature debonding. The embedment length of the steel connectors or anchoring of

longitudinal rebar must account for the anchorage failure mechanisms and the maximum roughness of the

interface. This parameter should also be controlled on site.

When compared to the reference specimens, detailing with rebars crossing the interface reached a

performance gain for each solution in terms of maximum shear stress at the interface. Rebar crossing the

interface with greater anchorage length resulted in a performance gain of more than double the shear

stress of the reference specimens. Even with insufficient anchorage length, the reinforcement crossing the

interface resulted in a performance gain of 60 % to 110 % for the shear stress at the interface. For the

solution with both steel connectors and longitudinal rebar anchored, the shear stress at the interface was

three times that of the reference specimens.

Provisions on the Model Code 2010 [19] for concrete-to-concrete interfaces, considering the proposed

coefficients, fit well with the experimental results for specimens with proper anchoring of the

reinforcement crossing the interface.


18

This work, by simplifying the horizontal shear stress problem to one direction, allowed for a better

understanding of the concrete-to-concrete interaction phenomenon on a tensile face, and will allow for the

development of two direction shear specimens, characteristic for punching strength on flat slabs.

Acknowledgements This work was supported by Fundação para a Ciência e Tecnologia – Ministério da Ciência,

Tecnologia e Ensino Superior through project EXPL/ECM–EST/1371/2013 and

grant SFRH/BD/89505/2012. Contributions of CONCREMAT SA and SIKA AG are also acknowledged

for the production of concrete specimens and providing the bonding agent for anchoring the



19

References

[1] Münger F, Wicke M, Randl N. Design of Shear Transfer in Concrete-Concrete Composite

Structures. IABSE Reports, 1997, p. 163–8.

[2] Randl N. Design recommendations for interface shear transfer in fib Model Code 2010. Struct

Concr 2013;14:230–41.

[3] Walraven JC, Reinhardt HW. Theory and experiments on the mechanical behavior of cracks in

plain and reinforced concrete subjected to shear loading. Concr. Mech., vol. 26, 1981, p. 65.

[4] Okamura H, Maekawa K. Nonlinear Analysis and Constitutive Models of Reinforced Concrete.

Tokyo: 1991.

[5] Vecchio FJ, Lai D. Crack Shear-Slip in Reinforced Concrete Elements. J Adv Concr Technol

2004;2:289–300.

[6] Birkeland PW, Birkeland HW. Connections in Precast Concrete Construction. J Am Concr

Institute, Proc 1966;63:345–67.

[7] Mattock AH, Hawkins NM. Shear Transfer in Reinforced Concrete−Recent Research. J Prestress

Concr Inst 1972;17:55–75.

[8] Loov RE, Patnaik AK. Horizontal Shear Strength of Composite Concrete Beams with Rough

Interface. PCI J 1994;39:48–69.

[9] Walraven JC. Aggregate interlock: A theoretical and experimental analysis. 1980.

[10] Randl N. Investigations on load transfer between old and new concrete at different surface

roughnesses. 1997.

[11] Silfwerbrand J, Beushausen H, Courard L. Bonded Cement Based Material Overlays for the

Repair, the Lining or the strengthening of slabs or pavements. Springer Netherlands 2011:51–79.

doi:10.1007/978-94-007-1239-3.

[12] Santos PMD, Júlio ENBS. Interface shear transfer on composite concrete members. ACI Struct J

2014;111:113–21. doi:10.14359.51686543.

[13] Júlio ENBS, Branco FB, Silva VD. Concrete-to-concrete bond strength. Influence of the

roughness of the substrate surface. Constr Build Mater 2004;18:675–81.

doi:10.1016/j.conbuildmat.2004.04.023.

[14] Silfwerbrand J. Improving Concrete Bond in Repaired Bridge Decks. Concr. Int., 1990, p. 61–6.

[15] Saucier F, Pigeon M. Durability of new-to-old concrete bonding. Proc. ACI Int. Conf. Eval.

Rehabil. Concr. Struct. Innov. Des., Hong Kong: 1991, p. 689–707.

[16] Abu-Tair AI, Rigden SR, Burley E. Testing the bond between repair materials and concrete

substrate. ACI Mater Journal1 1996;93:553–8.

[17] Bissonnette B, Courard L, Beushausen H, Fowler D, Trevino M, Vaysburd a. Recommendations

for the repair , the lining or the strengthening of concrete slabs or pavements with bonded cement-


20

based material overlays Recommandations pour la re ¸ age ou le ´ es en renforcement des dallages

industriels ou des chausse ´ ton par un rec. Mater Struct 2012;46:481–94. doi:10.1617/s11527-

012-9929-8.

[18] Transportation Research Board. Removing Concrete from Bridges. NCHRP Synth. 169,

Washington, D.C.: 1991, p. 42.

[19] Fédération Internationale du Béton (fib). Model Code 2010 - Final draft. vol. 1. Lausanne,

Switzerland: 2012.

[20] Branco FAB, Silva VD, Júlio ENBS. Concrete-to-concrete bond strength: influence of an epoxy-

based bonding agent on a roughened substrate surface. Mag Concr Res 2005;57:463–8.

doi:10.1680/macr.2005.57.8.463.

[21] Bissonnette B, Courard L, Fowler DW, Granju JL. Bonded Cement Based Material Overlays for

the Repair , the Lining or the Strengthening of Slabs or Pavements: State-of-the-Art Report of the

RILEM Technical Committee 193-RLS. 2011.

[22] Bissonnette B, Vaysburd AM, von Fay KF. Best Practices for Preparing Concrete Surfaces Prior

to Repairs and Overlays. 2012.

[23] RILEM Committee TC113. PC-5: Method of test for compressive strength of polymer concrete

and mortar. 1995.

[24] RILEM Committee TC113. PCM-8: Method of test for flexural strength and deflection of

polymer-modified mortar. 1995.

[25] Santos PMD, Júlio ENBS. Recommended improvements to current shear-friction provisions of

model code. 3rd fib Int. Congr., 2010.


21

4. Pulloff testing for characterizing the interface tensile resistance

Interaction between two concrete layers of different ages is well known and studied in civil engineering

applications, with provisions in several codes that guarantee structural integrity. Strengthening of existing

structures by adding a new concrete layer relies on the resisting mechanism that shall develop at the

interface with the existing structure. This mechanism is known and divided in three main components:

chemical adhesion and micro interlocking, inherent to base materials and contact surface conditions at the

time of casting; friction with aggregate interlocking and dowel action shall develop if reinforcement

crosses the interface. Chemical adhesion plays a major role in structural integrity for tensile loads vertical

to the interface, and therefore should be assessed. The components of the resisting mechanism as

described are characteristic of a relative shear deformation of the two layers at the interface. In a

strengthening situation with a concrete overlay the edge of the overlaid concrete is loaded in tension. This

is due to the phenomena occurring in these regions that causes peeling of the overlaid concrete as

depicted in FIG, which is emphasized if applied in a tensile region.

Figure 1 - Peeling moment from edge lifting phenomenon (adapted from [1]).

From FIG one can see the acting shear stresses at the overlay’s edge, contributing with mode II fracture

energy to tensile (mode I) fracture energy. Pull-off strength is then a well accepted parameter for

characterizing the tensile strength of the interface, as well as an all-purpose interface performance index

[2]. Although well accepted as a surface characterizing parameter, in [3] is referred that this parameter

alone doesn’t correctly assess the interface monolithic behaviour. Complementary shear bond tests are

then required to fully assess all directions of loading and phenomena at the interface. If no cracking of the

interface occurs, these effects can contribute positively for tensile strength on rough surfaces, since

concrete ridges of the substrate are loaded in compression and frictions forces resist loading imposed in

tension.

In [3] is stated that debonding in a strengthening situation occurs when a crack reaches the interface, and

the stress component normal to the interface exceeds the latter’s tensile strength. Bond between two

layers can be classified according to its effectiveness in fully or partial binding of the contact area. For

structural purposes rather than aesthetic, full bond between the two layers is a requisite since no

predictable behaviour can be assumed otherwise [4]. According to the authors, incomplete bond between

the two layers can cause a restraint on fully bonded zones, and cracking perpendicular to the latter,

according to FIG.

Figure 2 - Classification of bond: from complete to debonding.

This relates to one of the main characteristics of tensile strength in a rough interface, the effective surface

area, typically larger than geometric area, as depicted in FIG. This parameter is one of the most important


22

for bond capacity of concrete interfaces, according to [5] and [6]. Also, the effectiveness of mechanical

adhesion depends on the penetration of fresh concrete liquid matrix in the grooves of the existing surface,

which after hardening will help cohesion by interlocking the two layers.

Figure 3 - Contact area between two layers: geometric, true, and effective.

This interlocking phenomenon is relevant in tension for mechanical adhesion when effective vertical

anchorage in pores and voids exists, as depicted in FIG, along with chemical adhesion that plays a major

role in tensile resistance.

Figure 4 - Mechanical tensile and shear bond between two layers.

The contact area is defined in [7] with a transition zone between the two layers, which can also create a

layer of weakness since small aggregates can accumulate near the contact surface creating a wall effect as

depicted in FIG. The authors also refer to the behaviour between fresh and hardened concrete to be

similar to bond between aggregates and cement paste. In [8] this phenomenon was attested by shear bond

tests, where failure occurred in the overlaid concrete, and near the interface. A good correlation was

found between interface shear resistance and the overlaid concrete compressive strength, which is

particularly important for design purposes. In [9] the transition zone is also identified between the two

layers depicted in FIG, described as the zone where any interaction should occur that disturbs the stress

fields of the other layer.

Figure 5 - Transition zone between two layers ([7], [9]).

Another parameter that can rule the concrete overlay behaviour and resisting capacity is its thickness.

Previous studies suggest both likelihood of thin overlays to debond more easily than thicker ones, and no

dependence whatsoever on overlay thickness regarding bond strength [10]. Overlaid concrete thickness is

more likely to affect shear behaviour rather than tensile for larger thicknesses. In a strengthening situation


23

with bending stresses, thickness directly relates to the moment at the interface and irrespective tensile and

compressive stresses, according to FIG.

Figure 6 - Stresses acting in a finite element between cracks.

Although several methods exist [11], chipping with an electric hammer and high-pressure waterjet are the

most commonly used due to economy and availability for the first, and performance for the second. In [2]

caution is advised for the selection of surface preparation method according to it’s aggressiveness and the

substrate concrete strength class, with a lower limit of C30/37 class advised for the most aggressive

methods. Along with cracking of the overlaid concrete, microcracking of the existing surface can also

determine if good bond can develop at the interface. In a tensile situation, this is a critical detail since the

discontinuity due to cracking results in premature spalling of concrete and failure of the interface. In [12]

this problem is addressed when comparing several methods for existing surface preparation. Chipping

with an electric hammer causes microcracking of the substrate layer and is stated as one of the most

aggressive methods for surface preparation. Concrete removal with high-pressure waterjet does not

promote microcracking. The authors in [13] also refer to the depth where this phenomenon occurs,

between 3,0 mm and 10 mm, reducing interface performance. According to [12] the latter technique is the

only that removes concrete selectively, which is particularly important in deteriorated concrete scenarios.

Cleanliness of the interface is described in [4] as the most important parameter influencing interface

performance. The existing surface should be free of any contaminants, micro and macro particles, such as

dust and concrete from surface treatment, and moistened prior to concreting. Humidity is also of the

utmost importance in concrete-to-concrete interfaces since the porosity of the existing layer will draw

water from fresh concrete poured to the existing surface. In [14] several scenarios for humidity at the

existing surface were considered. The excessively dry and saturated surface resulted in poor bond,

whereas good bond resulted from apparently dry surface with saturated substrate. This falls in with test

results from [15], which recommends the same combination of dry surface and saturated substrate, with

limited fresh concrete water loss through porosity, as depicted in FIG.

Figure 7 – Substrate humidity: very dry, saturated surface, dry surface/saturated medium.

Other very important factor is the existing surface profile, which governs contact area and interlocking of

the two layers. Roughness is the ruling parameter for characterizing the surface capability for transferring

stresses in between layers. In [16] is stated that for concrete surfaces prepared with the aforementioned

methods, roughness and waviness parameters are enough for their assessment. The former parameter is

evaluated in terms of the average and total roughness of the surface profile, as depicted in FIG.


24

Figure 8 – Average roughness and mean peak-to-valley height for surface profile.

Several methods are considered in [17] for evaluating surface roughness, namely the sand patch method,

which is referred to as not suitable for very rough surfaces. Other methods include mechanical

profilometry [18] and laser profilometry [19], with the latter being the most accurate method available to

date. Code provisions in [20] state that the average roughness goes from smooth to very rough, with limit

values of 1.5 mm for the first, and 3.0 mm for the latter. Determination of the average roughness

parameter will reflect on design code provisions, where characterization of the surface roughness

quantifies adhesion and friction stresses acting on the interface. In [11] the qualitative character of surface

roughness assessment methods adopted in code provisions is mentioned, and stated that should be altered

to quantitative methods, like the ones already referred to in the MC 2010 [20].

Roughness alone doesn’t implicitly improve tensile strength. Although it improves the effective contact

area and macroscopic interlocking, adhesion comes from microscopic and sub-microscopic interlocking

of the two layers, both chemically and mechanically. In FIG an illustration of the same surface is depicted

as observed under different scales.

Figure 9 – Macroscopic, microscopic, and sub-microscopic surface roughness (adapted from [8]).

Overlaid concrete interacting with an existing structure will resist shrinkage stresses and loading when in

service. The edge-lifting phenomenon is particularly important when reinforcement is added to the new

concrete layer, with eccentricity of the reinforcement maximizing the peeling moment and tensile stresses

at the edge of the overlaid concrete. With this premise, the pull-off test is adequate for direct assessment

of tensile strength in this zone. This test is also one of the most popular tests for concrete-to-concrete

surface characterization due to its ability to be performed in situ [13].

Various tests have been developed through the years for assessing interface bond capacity, in tension, and

in shear [4]. Tensile strength can be assessed through pull-off testing of discrete specimens. This is the

most common test for tensile characterization of concrete-to-concrete interfaces is the pull-off test, with

reference of its use from 1984 in [21]. This type of test can be performed in the field by coring the

specimens in existing structures and the relatively light test equipment. Pull-off testing is a type of direct

tension test perpendicular to the interface, with inherent limitations pointed by several authors regarding


25

load eccentricity, damage during coring, and dispersion due to randomness of surface roughness and

tensile results ([14], [22], [23]). General setup for pull-off testing can be observed in FIG. High strength

epoxy glue is generally recommended for bonding the test dolly to concrete surface [22], with previous

cleaning of the laitance layer. Coring of the specimens should reach a depth of at least 25 mm through the

existing substrate, enough to assess the condition of existing concrete through tensile strength [4].

Figure 10 - Pull-off test setup (redo).

Failure modes for pull-off testing are divided in adhesive or cohesive, regarding whether debonding

occurs on an interface, dolly-concrete or concrete-concrete, or by tensile failure in any of the layers. An

illustration of each scenario is provided in FIG. Characteristics for pull-off test conditions and preparation

are provided in EN 1504-3 [24]. The European standard specifies minimum surface cohesion values for

structural repairs of 2,0 MPa and a minimum of five valid tests. Pull-off testing is directly addressed in

EN 1542 [25], regarding general aspects for this type of test, mainly dolly sizes, angle of force

application, and test procedure.

Figure 11 - Pull-off testing failure modes.

Viability for characterization of tensile performance through pull-off testing is validated according to [4],

since cracking in concrete occurs in a combination of modes I and II of fracture mechanics (tension, and

in-plane shear, respectively). The pull-off test characterizes directly tensile resistance for zones where

tensile stresses are predominant at the interface [26]. These zones are usually overlay boundaries, where

the edge-lifting phenomenon is present [1], due to lack of neighbouring elements restrained shrinkage and

equilibrium.

Motivation and testing An experimental investigation for tensile strength assessment was devised to allow for surface strength

characterization on several specimens, roughened prior to concreting a new layer on top of an existing

one.


26

Greater interaction between the two layers can be achieved by roughening the existing concrete surface

(substrate). In [12] is referred that this interaction can have an upper bound for tensile strength coherent

with values for sandblasted surfaces. Greater roughness can also expose deeper and larger aggregates,

along with more cement matrix area, which can help bonding the two layers. With these premises, surface

preparation was characterized for each specimen in order to account for changes due to roughness. The

technique chosen for preparing the surface was milling with an electric hammer and steel moil point.

Several tests were performed with only the laitance layer removed with a diamond grinder, in order to

assess surface preparation performance and effective area. Both resulting surfaces can be observed in

FIG.

Fig. 22 Two types of surface preparation: milling with steel moil point (left), and laitance layer removal

with diamond grinder (right).

Roughness was determined with the point measurement system illustrated in FIG, equivalent to the

mechanical stylus system described in [27]. This type of roughness assessment is referred in [16] as

adequate for concrete surfaces prepared with the more aggressive methods. This particular system allows

for forty equally spaced point measurements with a 50 mm TML CDP-50 displacement transducer, along

a 90 cm span. Several lines were evaluated for characterizing the surface roughness.

Fig. 23 Setup for surface roughness assessment.

Probe tip size was defined for the smallest machining that could be performed, since caution is advised in

[11], with a detailed illustration in FIG.


27

Fig. 24 Assessment for smaller grooves conditioned by the probe tip size (adapted from [28]) and probe

tip used.

This allowed for correct evaluation of the smaller grooves in the substrate surface. Although data for only

forty point measurements was acquired over a 1,0 m span, accuracy was acceptable for the macro

roughness of the evaluated specimens.

Prior to concreting the overlay, the substrate surface was cleared of any debris with pressurized air, and

wetted a few hours before casting until substrate was saturated. Approximately one hour before casting,

free-floating water was removed with pressurized air, for a dry surface/saturated substrate combination

upon casting of the overlay.

After casting the specimens to be tested, 90 days passed until pull-off tests were performed. Common

dolly sizes for pull-off testing are circular with 50 mm in diameter [25], or square with 100 mm side

length [29]. For concrete with an aggregate size up to 20 mm, an heterogeneous surface arises when

intercepting the interface, as shown in FIG.

Fig. 25 Difference in contact surface heterogeneity between different

areas.

Also, standard 50 mm dollies are relatively small when compared to the mean spacing between grooves

and ridges in the existing surface. In [30] is stated that rough surfaces should have at least a spacing of 40

mm between grooves or ridges. The intended surfaces to be classified as rough or very rough fall short of

this requisite with 50 mm dollies. A test setup for bigger dollies was then developed to allow for surface

characterization with a very rough profile and larger aggregates. A least heterogeneous contact surface

can be achieved with larger areas, as depicted in the previous figure. This balances the equilibrium

between aggregates of different sizes and cement matrix, allowing for better characterization of the

interface properties.

Several factors according to [22] determined the contact area for dollies used in this work. Expected

interface performance in tension results in higher loads for increased contact areas. This limits the test

setup, which needs to guarantee a stiff response for these loads until failure of the specimen. Also

stiffness of the metal dolly has to be guaranteed not to influence test results. A larger dolly would require

heavy core drilling for partial coring of the substrate layer below the test specimen. A square dolly allows

for a straight partial coring with a regular tool in the substrate. A parametric study regarding several dolly

dimensions was carried with ATENA 3D software, with test results presented in FIG.

Aggregates exposed

Cement surrounding

aggregates Cement matrix w/o agg.


28

Fig. 1 Pull-off test dolly detail.

For the same parametrization of the interface, considering 50 mm, 75 mm, 100 mm, and 150 mm square

dollies, the former yielded the largest tensile strength, and the latter the smallest tensile strength. After all

considerations, a dolly with a side length of 150 mm was determined, as depicted in FIG. This enables a

test setup that can be assembled by one person, and both test setup and steel dolly stiff enough that cause

no influence to test results for higher loads. Steel plates were welded symmetrically to the dolly’s axis for

strengthening the plate and preventing bending phenomena.

Fig. 2 Pull-off test dolly detail.

Test setup consisted on a steel plate with a side length of 300 mm and a thickness of 20 mm. Three holes

were threaded in the plate for the supporting 22 mm diameter threaded rods, allowing for the levelling of

the steel plate as required for this type of test. Detail can be observed in FIG.

Fig. 3 Reaction plate and threaded rod support for pull-off test setup.

Tensile forces were applied to the specimens through a threaded rod with 32 mm diameter fixed with a

hex nut welded to the steel dolly. A hollow plunger cylinder ENERPAC RCH-206 and a TML

CLC200KNA were used, both with 200 kN capacity. General test setup can be observed in FIG.

Hex nut


29

Fig. 4 Setup for pull-off testing.

The load requirements could be met with a less capacity cylinder, but the inner diameter requirement for

the plunger also resulted in the choice for this cylinder in particular. A high density elastomer with a

thickness of 10 mm was placed between the hydraulic jack and the steel plate to allow for correction of

any lack of verticality when preparing the specimens. It can be observed working for this purpose in FIG.

Loading was performed with an electronically controlled hydraulic pressure controller WALTER+BAI

PKNS19D at a rate of 0,005 MPa/s until failure. Loading history example and resulting failure can be

observed in FIG for one specimen.

Fig. 5 Elastomer for load alignment and load history example for one test.

Adhesive failure mode of the concrete-to-concrete interface was the general result for all tests. Exceptions

occurred through failure of the bonding agent at the steel dolly contact surface. Several measures were

taken to avoid this type of failure, namely the removal of the laitance layer through sandblasting, cutting

grooves, and cleaning of any particle of the concrete surface. Procedure and final aspect can be observed

in FIG. Bonding of steel dolly was carried out with an epoxy bonding agent HILTI HIT-RE500, whose

tensile bonding capacity largely surpasses the required for the expected loads, and consistency guaranteed

no influence to specimen base.

Fig. 6 Removal of laitance layer through sandblasting, groves for improved

dolly adhesion, and bonding of steel dolly.

High density

Elastomer

Hollow core

Plunger

Load cell


30

The adhesive failure at the concrete-to-concrete, characteristic for all tests, can be observed in FIG.

Debonding resulted clean through the interface, with the markings of the steel moil point still visible in

the substrate contact surface, and reversed in the debonded specimen surface. The failure mode

characteristic for all specimens can be a consequence of the more aggressive surface preparation method

chosen, according to test results in [31].

Fig. 7 Concrete-to-concrete adhesive failure of pull-off specimens.

Results and discussion Testing for surface characterization was carried out in several campaigns, coincidental with full size

specimens to be tested strengthened with concrete overlay that will be referred to as “OL”. This allowed

for characterization of several interfaces to be loaded in a strengthening situation and determination of

relevant surface parameters. Characteristics for materials of both layers in terms of compressive and

tensile strength, along with resulting tensile stresses for pull-off testing, were also accounted for.

A table is presented next with parameters for all specimens, regarding concrete compressive and tensile

strength for both layers, average and total surface roughness characteristics, and pull-off tensile strength.

Concrete compressive strength was considered for cylinder specimens.

Table 1 – Main parameters regarding surface strength and characteristics for all specimens.

The resulting pull-off strength varied from a minimum of 0.73 MPa to a maximum of 1.44 MPa average

stress. Distribution of the resulting stress for each specimen can be observed in FIG.

OL-1 OL-2 OL-3 OL-4 OL-5 OL-6 OL-7 OL-8 OL-9 OL-10

Substrate

(MPa)

fcs.m 34.8 26.4 25.6 32.8 44.4 45.1 43.8 36.2 29.9 32.5

fts.m 3.0 2.6 2.3 2.9 3.9 4.0 3.9 2.8 2.3 2.5

Overlay

(MPa)

fco.m 37.2 34.3 39.3 36.9 32.9 33.2 31.5 35.1 35.1 35.1

fto.m 2.6 3.2 2.8 2.9 3.1 3.1 3.0 2.9 2.9 2.9

Ra (mm) 3.8 3.3 3.1 3.6 2.9 2.5 3.0 2.8 2.9 2.9

Rt (mm) 9.2 7.4 8.3 9.1 7.9 8.2 8.4 9.9 10.6 10.2

σ (MPa) 0.82 1.04 0.82 0.85 1.34 1.22 1.44 1.36 0.73 1.16

SD (MPa) 0.18 0.19 0.22 0.31 0.25 0.21 0.12 0.12 0.13 0.15

COV (-) 0.22 0.18 0.26 0.36 0.19 0.17 0.08 0.09 0.18 0.13

Grooves from the surface

preparation phase


31

Fig. 8 Pull-off strength for all specimens tested.

Accounting for concrete compressive strength for both layers, a positive correlation factor of 0.73 was

found when the substrate concrete was compared with the pull-off strength for all specimens. A negative

correlation factor of -0.77 was found when the latter was compared with the overlaid concrete

compressive strength, which indicates an inverse relationship between the two. Both results can be

graphically observed in FIG.

Fig. 9 Relationship between pull-off strength and concrete compressive strength for both

layers.

Regarding concrete tensile strength, pull-off test results yielded a positive correlation factor of 0.68 for

substrate concrete tensile strength, and 0.53 for the tensile strength of the overlaid concrete. Both can be

observed in FIG.

Fig. 10 Relationship between pull-off strength and concrete tensile strength for both layers.

To assess the effectiveness of surface preparation (SP), a simpler technique with a diamond grinder (DG)

was also tested. This comparison was carried out only for OL-5, OL-6, and OL-7 specimens. Results are

presented in TAB, and compared with the previous surface preparation technique with steel moil point

(SMP).

0,0

0,4

0,8

1,2

1,6

2,0

1 2 3 4 5 6 7 8 9 10P

ull

off

str

ength

[N

/mm

2]

0,0

0,5

1,0

1,5

2,0

2,5

0

10

20

30

40

50

Pu

ll-o

ff s

tren

gth

[N

/mm

2]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

fcs,m AVG (MPa)

0,0

0,5

1,0

1,5

2,0

2,5

0

10

20

30

40

50

Pu

ll-o

ff s

tren

gth

[N

/mm

2]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

fco,m AVG (MPa)

0,0

0,5

1,0

1,5

2,0

2,5

0,0

1,0

2,0

3,0

4,0

5,0

Pu

ll-o

ff s

tren

gth

[N

/mm

2]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

fts,m AVG (MPa)

0,0

0,5

1,0

1,5

2,0

2,5

0,0

1,0

2,0

3,0

4,0

5,0

Pu

ll-o

ff s

tren

gth

[N

/mm

2]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

fto,m AVG (MPa)

σPull-off σPull-off



32

Table 2 – Comparison between surface preparation techniques.

OL-5 OL-6 OL-7

Specimen SMP DG SMP DG SMP DG

1 1.03 0.99 1.04 0.70 1.44 0.80

2 1.42 0.85 1.20 0.55 1.51 0.66

3 1.83 1.06 0.98 0.84 1.46 1.14

4 1.16 1.02 1.62 0.35 1.20 1.02

5 1.25 0.91 1.35 0.71 1.57 1.26

6 1.38 1.13 1.14 0.49 1.43 1.41

σ (MPa) 1.34 0.99 1.22 0.61 1.44 1.05

SD (MPa) 0.25 0.09 0.21 0.16 0.12 0.26

COV (-) 0.19 0.10 0.17 0.26 0.08 0.25

SMP/DG (-) 1.35 2.01 1.37

Results show for a more aggressive surface preparation technique an increment of approximately one

third of the tensile load for the OL-5 and OL-7 specimens, and double the tensile strength for OL-6.

Graphically, a relationship between the two scenarios can be observed in FIG. Pull-off specimens 1 to 6

of each surface preparation were placed next to each other for minimum variation of substrate concrete

characteristics across its surface.

Fig. 11 Impact of surface preparation on interface tensile strength.

This was a direct assessment for the impact of surface preparation on tensile strength between two

concrete layers. Despite being the specimen that most benefitted from surface preparation, according to

Table 1 the OL-5 specimen was also the least rough of the three tested. This suggests an upper bound for

surface roughness in tension, also according to a conclusion referred in [12]. Since for very rough

surfaces there is no significant increment to the effective area of contact, this can justify the impact of a

smaller average roughness could result in higher tensile strength.

Values for average surface roughness fell within normal values for rough to very rough surfaces as

defined by several authors and design codes. The total roughness of the profile was also computed. This

parameter, which states the relationship between highest peaks and lowest valleys, can attest the

interlocking capacity of the two layers for larger relative displacements. Comparing these values with

pull-off strength results yields a relationship between these two parameters, as observed in FIG.

0,0

0,5

1,0

1,5

2,0

1 2 3 4 5 6

Pu

llo

ff s

tren

gth

[N

/mm

2]

Specimen

OL-4

1 2 3 4 5 6

Specimen

OL-5

1 2 3 4 5 6

Specimen

OL-6

w/ SP

w/o SP


33

Fig. 12 Relationship between substrate concrete compressive and tensile strengths, and average

surface roughness.

It should be noted that surface roughness is a user dependent variable when preparing the surface for

casting the new layer, and not exclusively material dependent. In FIG both average and total roughness

are compared with concrete compressive and tensile strengths for the substrate layer.

Fig. 13 Relationship between substrate concrete compressive and tensile strengths, and average

surface roughness.

Surface classification through these values according to [ , , ] can’t be absolute, since surfaces with the

same average roughness can present a very different profile. Other parameters mentioned in PARAGraph

were then evaluated based on roughness assessment results, and are presented in TAB.

Conclusions Roughness assessment and pull-off tests were carried out to characterize the contact area between

concrete layers cast at different times. These methods allowed characterization of the substrate surface

prior to casting the concrete overlay, and correlation to test results.

Surface roughness was assessed through a point measurement setup which evaluates linear spans of a

surface area. A surface characterization of rough or very rough could be performed due to the resulting

surface profiles, with average roughness varying from 2.5 (OL-5) to 3.8 (OL-1). The adopted system

worked with enough precision for this type of macro roughness.

0

2

4

6

8

10

0,0

0,5

1,0

1,5

2,0

Aver

age

rou

gh

jnes

s [m

m]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

AVG (MPa) Ra (mm)

0

10

20

30

40

0,0

0,5

1,0

1,5

2,0

Aver

age

rou

gh

nes

s[m

m]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

AVG (MPa) Rt (mm)

0

2

4

6

8

10

0

10

20

30

40

50A

ver

age

rou

gh

nes

s [m

m]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

fcs,m Ra (mm)

0

2

4

6

8

10

0,0

1,0

2,0

3,0

4,0

5,0

Aver

age

rou

gh

nes

s [m

m]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

fts,m Ra (mm)

0

10

20

30

40

0

10

20

30

40

50

Tota

l ro

ugh

nes

s [m

m]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

fcs,m Rt (mm)

0

10

20

30

40

0,0

1,0

2,0

3,0

4,0

5,0

Tota

l ro

ugh

nes

s [m

m]

Con

cret

e te

nsi

le s

tren

gth

[N/m

m2]

fts,m Rt (mm)



34

Pull-off testing was performed with a setup devised specifically for this work, which worked well and

allowed for assessing the interface tensile strength. The larger dollies allowed for a more homogeneous

contact surface for the specimens, and better results with less variability for each specimen.

Tensile strength for the interface varied from 0.73 MPa to 1.44 MPa, with the latter representing almost

double the tensile strength of the former. This variability can be explained since materials were different

for each set of specimens (OL-1 to OL-10) and surface preparation, which can’t be prescribed exactly for

each specimen. Curing and relationship between time of casting and testing were also different for each

set of tests.

Correlation…


35

References [1] Granju J-L. Debonding of Thin Cement-Based Overlays. J Mater Civ Eng 2001;13:114–20.

doi:10.1061/(ASCE)0899-1561(2001)13:2(114).

[2] Courard L, Piotrowski T, Garbacz A. Near-to-surface properties affecting bond strength in

concrete repair. Cem Concr Compos 2014;46:73–80. doi:10.1016/j.cemconcomp.2013.11.005.

[3] Perez F, Morency M, Bissonnette B. Correlation between the roughness of the substrate surface

and the debonding risk. Civ Eng 2009:949–56.

[4] Silfwerbrand J, Beushausen H, Courard L. Bonded Cement Based Material Overlays for the

Repair, the Lining or the strengthening of slabs or pavements. Springer Netherlands 2011:51–79.

doi:10.1007/978-94-007-1239-3.

[5] Santos PMD, Júlio ENBS, Silva VD. Correlation between concrete-to-concrete bond strength and

the roughness of the substrate surface. Constr Build Mater 2007;21:1688–95.


[6] Talbot C, Pigeon M, Beaupre D, Morgan DR. Influence of surface preparation on long-term

bonding of shotcrete. ACI Mater J 1994;91:560–6.

[7] Saucier F, Pigeon M. Durability of new-to-old concrete bonding. Proc. ACI Int. Conf. Eval.

Rehabil. Concr. Struct. Innov. Des., Hong Kong: 1991, p. 689–707.

[8] Beushausen H-D. Long-term performance of bonded concrete overlays subjected to differential

shrinkage. 2005.

[9] Emmons PH, Vaysburd AM. Performance Criteria for Concrete 1995.

[10] Bissonnette B, Courard L, Beushausen H, Fowler D, Trevino M, Vaysburd a. Recommendations

for the repair , the lining or the strengthening of concrete slabs or pavements with bonded cement-based

material overlays Recommandations pour la re ¸ age ou le ´ es en renforcement des dallages industriels ou

des chausse ´ ton par un rec. Mater Struct 2012;46:481–94. doi:10.1617/s11527-012-9929-8.

[11] Santos PMD, Júlio ENBS. A state-of-the-art review on roughness quantification methods for

concrete surfaces. Constr Build Mater 2013;38:912–23. doi:10.1016/j.conbuildmat.2012.09.045.

[12] Silfwerbrand J. Improving Concrete Bond in Repaired Bridge Decks. Concr. Int., 1990, p. 61–6.

[13] Júlio ENBS, Branco FB, Silva VD. Concrete-to-concrete bond strength. Influence of the

roughness of the substrate surface. Constr Build Mater 2004;18:675–81.


[14] CLELAND DJ, LONG AE. THE PULL-OFF TEST FOR CONCRETE PATCH REPAIRS. Proc.

Inst. Civ. Eng. Struct. Build., vol. 122, n.d.

[15] Emmons PH. Concrete Repair and Maintenance Illustrated: Problem Analysis; Repair Strategy;

Techniques. 1992.

[16] Garbacz a., Courard L, Bissonnette B. A surface engineering approach applicable to concrete

repair engineering. Bull Polish Acad Sci Tech Sci 2013;61. doi:10.2478/bpasts-2013-0006.

[17] Merl RN. Best Practices for Preparing Concrete Surfaces Prior to Repairs and Overlays 2012.

[18] Abu-Tair AI, Rigden SR, Burley E. Testing the bond between repair materials and concrete

substrate. ACI Mater Journal1 1996;93:553–8.

[19] Júlio ENBS, Santos PMD. Development of a laser roughness analyser to predict in situ the bond

strength of concrete-to-concrete interfaces. Mag Concr Res 2008;60:329–37.

doi:10.1680/macr.2007.00024.

[20] Fib. Model code 2010 Vol.1 2012:247–78.


36

[21] Long AE, Murray AM. The “Pul l-Off” Partially Destructive Test for Concrete. Spec Publ

1984;82:327–50.

[22] Vaysburd A, McDonald J. An evaluation of equipment and procedures for tensile bond testing of

concrete repairs. Tech Rep REMR-CS-61, US Army Corps Eng Washington, DC 1999.

[23] Delatte by NJ, Jr., Wade DM, Fowler DW. Laboratory and Field Testing of Concrete

BondDevelopment for Expedited Bonded Concrete Overlays. Mater J 2000;97:272–80.

[24] EN 1504-3:2005 - Products and systems for the protection and repair of concrete structures -

Definitions, requirements, quality control and evaluation of conformity - Part 3: Structural and non-

structural repair n.d.

http://standards.cen.eu/dyn/www/f?p=204:110:0::::FSP_PROJECT,FSP_ORG_ID:24608,6087&cs=1E52

D70FC476EEADAF818E66E11242299 (accessed October 31, 2015).

[25] BS EN 1542:1999. Products and systems for the protection and repair of concrete structures. Test

methods. Measurement of bond strength by pull-off. BSI; 1999.

[26] Carter P, Gurjar S, Wong J. Debonding of Highway Bridge Deck Overlays. Concr Int

2002;24:51–8.

[27] Whitehouse D. Surfaces and Their Measurement. Elsevier; 2002. doi:10.1016/B978-190399601-

0/50017-7.

[28] Bhushan B. Surface Roughness Analysis and Measurement Techniques. Mod Tribol Handb 2001.

[29] NBR 13528:2010. Render made of inorganic mortars applied on walls - Determination of bond

tensile bond strength. 2010.

[30] Eurocode 2 - Design of concrete structures Part 1-2: General rules Structural fire design

2010:109.

[31] Courard L, Sustercic J. Surface properties of concrete and criteria for adhesion of repair systems

Andrzej Garbacz and Tomek Piotrowski n.d.


37

5. Numerical analysis for unidirectional flat slabs strengthened with BCO

Introduction Adding a new concrete layer to an existing structure has been widely studied since it is an economical

way for structural strengthening with a well known material in civil engineering. The resisting mechanism

that bond the two concrete layers has been analysed and is widely accepted divided in three components [

]: adhesion, friction and aggregate interlocking, and dowel action of the reinforcement crossing the

interface. These three components can act together or individually, depending on load and the interface

detailing. The first component is more suitable for serviceability loads, and the other two are

characteristic for ultimate limit state loads. The latter one is exclusive to larger structural deformation

since for an initially monolithic interface, it’s the last component to resist interface loads.

Figure 12 - Components of the resisting mechanism for shear friction

This load transfer is divided orthogonally in vertical stresses that can cause failure by premature

debonding of the two layers, and horizontal shear stresses that ductilize the interface behaviour. To better

understand this stress distribution, a non-linear numerical model was analysed to correctly design the

connection between two concrete layers cast at different times. Nonlinear analytical software for

structural concrete can predict structural failure by a constitutive model for concrete with several material

parameters defined according to experimental data. These are usually based on a predictor-corrector

method for the analysis, mainly the Modified Newton Raphson method with tangent predictor and line

search [ ].

The lack of some experimental analysis to determine all the necessary parameters allows for definition

through theoretical or mathematical models based on empirical data in the bibliography. An important

aspect for modelling structures strengthened with a new concrete layer that utilizes empirical behaviour

models is the interface between the two layers. Several authors propose behaviour models to characterize

the resisting capacity of concrete-to-concrete interface based on the shear friction theory ([ ] ~ [ ]). These

are based on several parameters for the interface, mainly roughness of the substrate surface, compressive

strength of added concrete, and if reinforcement crossing the interface exists. These parameters rule the

behaviour of the interface with softening of stresses as the load increases and the interface crack opens.

The main difference is the reinforcement crossing the interface which can alter failure from brittle to

more ductile and the relationship between stress and relative displacements at the interface.


38

Concrete-to-concrete interface This will be the main topic to be analysed in this work since it governs global behaviour of structures

strengthened with an overlaid concrete layer. When reinforcement crossing the interface is prescribed, the

ductile behaviour of the interface results in relative sliding of both layers due to strain in the structure.

This relative slip then relates to the shear stress at the interface, as considered by several models [# - #],

which depends on the roughness profile, crack opening, and resisting strength of concrete. These three

parameters determine how much contact exists between the two layers, and if concrete ridges crush due to

relative slip. Several authors have proposed design methods to assess the behavior of the interface

between two different concretes, ranging from the shear-friction theory proposed by [ ], to the modified

shear friction theory proposed by [ ]. Design codes state the use of shear-friction based expressions to

calculate the resisting strength of the interface. The design recommendation presented in the Model Code

2010 [ ], which derived from the work of Randl [ ] on extended shear-friction theory, accounts for all

major parameters of a concrete-to-concrete interface. This design model is divided in two major scenarios

for the interface, rigid and non-rigid, depending on whether the interface crack opening is smaller or

greater than 0,05 mm. This allows for a better adjustment to real interface detailing, regarding the

existence of reinforcement crossing the interface, which in turn defines the brittle or ductile behavior.

Then the model is further divided in three mechanisms, adhesion, friction and aggregate interlocking, and

dowel action of the reinforcement.

Although this design model allows for determining the interface resisting strength in shear, there’s no

clear relationship with the relative slip of the two layers. Walraven-Reinhardt (1981) [ ], Okamura-

Maekawa (1991) [ ], and Lai-Vecchio (2001) [ ], all have proposed models that characterize the

relationship of the relative slip of concrete to concrete interfaces with shear stress, as described in Tab. 1.

Tab. 1 - Relative slip of concrete to concrete interfaces according to several models

Walraven-Reinhardt

(1981) 𝜹 =

𝒗𝒄𝒊 + 𝒗𝒄𝒐𝟏, 𝟖𝒘−𝟎,𝟖 + (𝟎, 𝟐𝟑𝟒𝒘−𝟎,𝟕𝟎𝟕 − 𝟎, 𝟐𝟎). 𝒇𝒄𝒄

Okamura-Maekawa

(1991) 𝜹 = 𝒘.√

𝝍

𝟏 −𝝍

Lai-Vecchio

(2001) 𝜹 =

𝟎, 𝟓𝒗𝒄,𝒎á𝒙 + 𝒗𝒄𝒐𝟏, 𝟖𝒘−𝟎,𝟖 + (𝟎, 𝟐𝟑𝟒𝒘−𝟎,𝟕𝟎𝟕 − 𝟎, 𝟐𝟎). 𝒇𝒄𝒄

. √𝝍

𝟏 − 𝝍

Where 𝒗𝒄𝒊 is the shear stress at the interface and 𝒘 the interface crack opening, which relates to the

separation of the surfaces. The direct relationship between these values when evaluating the interface

behavior because separation of the surfaces implies loss of contact, and thus no stresses can be

transferred. Surface separation, as given by the second model, is the relationship between the principal

strain in a given element at interface level, according to Eq. (1).

𝒘 = 𝜺𝟏. 𝒔 (1)

These models allow for estimating shear stress at the interface level, since it is the most difficult value to

measure in experimental models. Common to all models is also the dependence on concrete compressive

strength, which arises from friction at the interface. This is directly related to crushing of the protuberant

concrete ridges due to interface rough profile. Relative slip and surface separation can be actively

measured in the latter, allowing for some empirical calibration of the previous models. Note that these

models don’t account for rebar crossing the interface, which alters both stiffness and the behavior of the


39

interface, allowing for bigger vertical and consequent horizontal relative displacements. These factors

weight in on the softening of the relationship between stress and relative displacement at the interface.

Several authors propose various models to describe this relationship [# - #], from linear to multilinear [#,

#], to exponential or hyperbolic functions [#, #].

Softening of the relationship between stress and relative displacements is also dependent on

reinforcement crossing the interface, thus being an important parameter for choosing the softening model.

When no reinforcement crosses the crack, a linear or multilinear stress-displacement curve like the one in

FIG can be applied.

Fig. 1 – Fracture energy for when no rebar is crossing the interface.

One can observe in the figure the first elastic stage until tension limit stress is reached and surface

separation occurs. From that moment, bond is lost, and a friction resisting mechanism due to relative

sliding of the surfaces starts to develop. The kink point is referenced in the bibliography and determined

according to empirical data collected from various tests [# -#]. When a minimum amount of

reinforcement crosses the crack, a progressive softening occurs, with the exponential and power curves

best fitting the relationship between stress and relative displacement. This is characterized by the

subsequent opening of new cracks and crushing of the concrete ridges, along with bending of the

reinforcement as depicted in FIG.

Fig. 2 – Softening of stresses for when a minimum amount of rebar crosses the interface.

It is important for a nonlinear modelling software that includes the definition of an interface between two

materials to allow a softening law for that interface. FE modelling of the concrete-to-concrete interface

can be performed with discrete or zero-thickness elements. The latter elements are composed as a set of

parameters that characterize the stiffness and resisting strength for the interface. The constitutive model


40

for a general three-dimensional case is given in terms of normal and tangential stresses, and irrespective

relative displacements by:

{𝝉𝟏𝝉𝟐𝝈} = [

𝒌𝒕𝒕 𝟎 𝟎𝟎 𝒌𝒕𝒕 𝟎𝟎 𝟎 𝒌𝒏𝒏

] . {𝚫𝒗𝟏𝚫𝒗𝟐𝚫𝐮

} (2)

For a two-dimensional analysis the second row and column can be omitted. 𝒌𝒕𝒕 and 𝒌𝒏𝒏 denote the initial

elastic shear and normal stiffness respectively. It is recommended to estimate the stiffness value using the

following:

𝒌𝒏𝒏 =

𝑬

𝒕 (3)

𝒌𝒕𝒕 =

𝑮

𝒕=

𝒌𝒏𝒏𝟐. (𝟏 + 𝝊)

(4)

Interface width 𝒕 is a relative term when evaluating rough concrete interfaces because no discrete

interface can be identified, hence no direct value cane be measured. The software authors recommend an

approximate measure for this parameter, regarding the lowest young modulus and irrespective FE

dimensions [ ]. The initial failure surface corresponds to the Mohr-Coulomb condition with tension cut

off, depicted in FIG.

Fig. 3 - Failure surface at the interface

This surface is governed by Eq. () until tensile limit stress is reached, then shear stress becomes null

because no contact exists between the surfaces.

|𝝉| ≤ 𝒄 + 𝝈.𝝋 for 𝝈 ≤ 𝒇𝒕 (5)

𝝉 = 𝟎 for 𝝈 ≥ 𝒇𝒕 (6)

After stresses violate this condition, the failure surface collapses to a residual surface. This is due to the

larger displacements that no longer include the material cohesion or tensile strength, accounting only for

friction between the two layers. The tensile and shear softening are defined based on the fracture energy

associated with each mode.

3D Specimen modelling For analysing the stress distribution in the experimental specimens tested, numerical models to scale were

defined in the ATENA software, with the different detailing of the interface. The concrete material as

defined by the software is based on the SBETA constitutive model. This model mainly includes nonlinear


41

behaviour in compression, softening and hardening of the material, cracking in tension, tension stiffening,

and fixed and rotated crack modelling [ ]. Although this material can include rebar in a quasi-perfect bond

connection, these elements were modelled by linear elements embedded in concrete, with a bond-slip law

to govern the steel-concrete interface. The specimens double symmetry allowed to model only one quarter

of each specimen, thus saving time in calculating symmetrical coordinates with equal conditions, as

shown in FIG. A full model was also tested for reference (FIG).

Fig. 1 Quarter (left) and full (right) numerical models of experimental specimens.

Since FE modelling is sensitive to geometry of the elements to be modelled, there was a concern with

guaranteeing the least irregular geometries and the most full contacts of the adjacent elements surfaces.

This premise lead to a great number of contact surfaces, but also allowed to generate a regular mesh with

quadrilateral elements, despite the hexagonal shape of the rebar modelled with 3D finite elements (FIG).

Fig. 2 Contacts for modelled specimens with rebar anchored (left) and detail (right) .

The software developers refer the inability of linear elements to perform under bending or shear, thus

recommending several equivalent solutions to simulate this behavior [ ]. One solution is to model the

circular shape of the steel bars through 3D finite elements with a square or hexagonal shape, with

equivalent area and perimeter. For this matter, hexagonal shaped elements for reinforcement crossing the

interface was defined for the anchoring of longitudinal rebars, and square shaped elements for the steel

connectors (FIG). The option for square elements instead of hexagonal was due to the lower threshold of

the FE mesh, unable to generate such small elements.


42

Fig. 3 Detail of contacts for numerical models of STC experimental specimens.

The monitoring points assigned were the same as instrumented in the experimental test specimens, in

order to assess and compare each measurement. An example can be observed in FIG for comparison of

both instrumented environment.

Fig. 4 Monitoring points coincidental with experimental specimens.

The interface itself as defined by the software consisted on zero thickness elements, defined geometrically

by the contact area between the two layers for each specimen. The input parameters for a 3D interface on

ATENA software are stiffness, normal and tangential to surface, tensile strength normal to surface,

cohesion, and a friction coefficient. The tensile strength was taken directly from experimental pulloff

tests, presented in the TABLE below.

Table 3 – Main parameters regarding surface strength and characteristics for all specimens [REF].

Determination of the friction coefficient was carried out according to the proposal by [JULIO], referred in

[ ] for an average roughness of 3,0 mm as 1,4. Based on the Mohr Coulomb model illustrated in FIG a

cohesion was calculated for this friction coefficient. A softening of the stresses can also be defined by the

OL-1 OL-2 OL-3 OL-4 OL-5 OL-6 OL-7 OL-8 OL-9 OL-10

Substrate

(MPa)

fcs.m 34.8 26.4 25.6 32.8 44.4 45.1 43.8 36.2 29.9 32.5

fts.m 3.0 2.6 2.3 2.9 3.9 4.0 3.9 2.8 2.3 2.5

Overlay

(MPa)

fco.m 37.2 34.3 39.3 36.9 32.9 33.2 31.5 35.1 35.1 35.1

fto.m 2.6 3.2 2.8 2.9 3.1 3.1 3.0 2.9 2.9 2.9

σ (MPa) 0.82 1.04 0.82 0.85 1.34 1.22 1.44 1.36 0.73 1.16

SD (MPa) 0.18 0.19 0.22 0.31 0.25 0.21 0.12 0.12 0.13 0.15

COV (-) 0.22 0.18 0.26 0.36 0.19 0.17 0.08 0.09 0.18 0.13

Monitoring points


43

user. Determination of the stress softening law was chosen according to relative slip measurements during

experimental testing, as shown in FIG.

Fig. 5 Example for relative displacement at the end of the overlay (STANC detailing).

The model that best described the relationship between stress and relative displacement at the interface

was Reinhardt’s exponential curve [ ]. Stress for each relative displacement was calculated with EQ,

where n is a fitting parameter, calculated according to EQ, where GF,i is the interface fracture energy.

𝝈(𝒘) = 𝒇𝒕 {𝟏 − (

𝒘

𝒘𝒄)𝒏

} (7)

𝑮𝑭,𝒊 = ∫𝝈(𝒘) (8)

The latter parameter was considered proportional to the lower strength concrete fracture energy for each

specimen, divided by interface detailing and according to EQ, which derives from the Model Code 1990 [

]. The interface fracture energy is then calculated considering the weight of the tensile strength of the

interface and the weakest concrete irrespective fracture energy.

𝑮𝑭,𝑾𝒌𝑪𝒐𝒏𝒄 = 𝑮𝑭𝟎 (

𝒇𝒄𝒎𝒇𝒄𝒎𝟎

)𝟎,𝟕

(9)

𝑮𝑭,𝒊 =

𝒇𝒕,𝒊𝒇𝒕,𝑪𝒐𝒏𝒄

. 𝑮𝑭,𝑾𝒌𝑪𝒐𝒏𝒄 (10)

To calculate interface stiffness, a variation of the concrete stress-strain relationship was used, considering

the very small displacement needed for concrete to crack. Therefore, a complete stress-crack opening

relationship is proposed in FIG.

0

40

80

120

160

200

240

0,0 2,0 4,0 6,0

Lo

ad [

kN

]

Relative displacement [mm]

Relative displacement at overlay end

R4 vertical

R5 0,50m

R6 0,25m

σ

δ

Knn

1

𝑮𝑭,𝒊


44

Fig. 6 “Stress-Crack opening” relationship for the interface, pre and post cracking.

The positive slope, or interface stiffness KNN (or KTT), is governed by the interface fracture energy needed

for cracking to occur, and calculated according to EQ.

𝑲𝑵𝑵 =

𝒇𝒕,𝒊𝟐

𝟐. 𝑮𝑭,𝒊 (11)

For simplification, it is recommended by the software developers to consider KTT with the same value of

KNN. This procedure was adopted in this work with good results.

Results The individual material elements were also modelled to correctly characterize the behaviour of the

strengthened specimens. Therefore, simplified models were tested for concrete strength, both tensile and

compressive, rebar strength, in tension and pullout, and pulloff strength between two concrete layers.

For characterizing concrete strength in compression, a simple cube loaded in compression was modelled,

and for tensile strength, a prismatic specimen with twice the length of the previous cube was modelled,

loaded in tension. Simple one degree of freedom boundary conditions were assigned. Both FE models can

be observed in FIG, with the irrespective resulting curves. There was a good correlation with the

experimental results, with differences smaller than 5%.

Fig. 7 FE modelling of concrete tensile strength.

Fig. 8 FE modelling of concrete compressive strength.

Cracks


45

For the tensile strength of rebars a single linear element loaded in tension was modelled. This allowed to

identify exactly the strain for which yielding of the rebar occurred. There was also a good correlation with

the experimental characterization tests for the rebar illustrated in FIG.

Pullout of the rebar embedded in the substrate layer was modelled for a single rebar embedded in a larger

concrete substrate, as shown in FIG, so stresses can dissipate and no boundary conditions influence the

anchorage behaviour.

Fig. 9 FE modelling of rebar pullout strength when embedded with cement grout.

Comparison with experimental data shows good correlation, with differences below 5%, as shown in

TAB. This also allowed to analyse local crack dispersion in the substrate layer, illustrated in FIG, due to

anchorage pullout effects.

Pulloff strength between the two concrete layers was assessed by modelling a single concrete cube

bonded to a larger concrete substrate in order to impose volumetric discontinuity in the elements, as

shown in FIG. This causes a stress concentration in these areas, in accordance with the real specimen’s

volumetric discontinuities. The contact interface between the two elements was defined as described in

PARG for the reference specimens without reinforcement crossing the interface. Results regarding

concrete and interface cracking can be observed in FIG.

Fig. 10 FE modelling of overlaid concrete interfacial tensile strength.

After numerical characterization of the materials, the quarter models of the experimental specimens were

tested. The load step was defined as an increment of 0,1 mm at midspan, with good stability for Newton-

Raphson solution method of the FE analysis. Loading of the reference specimens can be divided in two

stages, until debonding of the two layers, and loading of the substrate layer until flexural failure. The

phenomena regarding interfacial debonding were identified in the analysis, with the plateau in the load-

deformation curve identifying this stage in the loading. In the figure below, the numerical and


46

experimental load-deformation curves show a good correlation, with both debonding and flexural failure

characterizing the behaviour for the reference specimens. One can see in FIG the modelling performed

considering only the substrate layer for comparison with strengthened specimens. Since there were no

experimental specimens with only substrate, the added value for these FE models is recognized.

Fig. 11 Correlation between numerical and experimental results for REF specimens.

In FIG a comparison between the two is presented, comparing the crack pattern and similar macro

cracking, with surface cracks coincidental for both numerical models and experimental specimens.

Fig. 12 Cracking for the strengthened cross section (left) and substrate only (right).

Partial debonding of the contact area in the numerical model shows good correlation with experimental

specimens observable behaviour, and correct functioning of the software’s interface definition. The

debonding phenomenon leads to a halt in crack formation, with closing of hairline cracks upon

debonding, hence no apparent macro cracking of the overlaid concrete layer, as shown in FIG.

Fig. 13 Debonding of the overlaid concrete, experimental (left) and numerical (right).

Analysing the stress distribution for the interface alone, one can identify a pattern for normal stresses until

debonding. The interface is loaded in tension at midspan and in the edge of the strengthened area, and

compressive stresses in between. These compressive stresses arise from the strut that forms from the

loaded area at midspan to the top of the slab. As the load increases, so does the global deformation of the


47

strengthened slab, and the consequently larger relative displacements at the interface lead to the

debonding of the two layers. This evolution can be observed in FIG. This debonding starts at the edge of

the strengthened area, leading to a reduction in stiffness observable in FIG, and evolves to midspan. Full

debonding occurs when not enough contact area exists between the two layers, with stiffness decreasing

to almost zero.

The second stage of the loading then starts by loading the substrate layer, until flexural failure by yielding

of the longitudinal rebar occurs. This stage can be observed in FIG, with an apparent stiffness until

failure.

For the modelling of specimens with reinforcement crossing the interface a good correlation was also

attained, with the altered behaviour when compared to the reference ones. Numerical models for both

anchorage depths adjusted accordingly to the experimental results. A comparison between the numerical

and experimental results is shown in FIG for the ANC and STC detailings for the 50 mm anchorage

length, and in FIG for the 70 mm anchorage depth. The comparison for the STANC detailing, which

combines all detailing parameters, is shown in FIG.

Fig. 14 Correlation between numerical and experimental results for the ANC specimens.

Fig. 15 Correlation between numerical and experimental results for the STC specimens.

Debonding of the overlaid concrete was not so apparent when plotting the load-deformation relationship,

observable only for the ANC detailing with 50 mm anchorage depth. These were the only specimens with

reinforcement crossing the interface that fully debonded in the experimental programme [ ]. The

specimens showed a more monolithic behaviour, with both layers bonded until failure. The crack patterns

for each detailing of the interface can be observed in FIG, showing good correlation to surface macro-

cracking observed for the experimental specimens.

-

40

80

120

160

200

240

- 5 10 15 20 25 30 35

Lo

ad

at

mid

spa

n [

kN

]


0

40

80

120

160

200

240

0 5 10 15 20 25 30 35

Lo

ad

at

mid

spa

n [

kN

]


ATENA

ATENA


48

Fig. 16 Correlation between numerical and experimental results for REF specimens.

When observing the resulting deformed structure in the software’s post processor, the resulting shear and

bending deformation of the reinforcement crossing the interface becomes apparent, as illustrated in FIG.

This behaviour shows evidence of the dowel action of rebars due to relative displacements of both layers

for tension and shear.

Fig. 17 Exaggerated deformed shape for the ANC specimens to show dowel action of the


Analysing the stress distribution at the interface shows a front of tensile stresses evolving from the

overlay’s edge to midspan, causing the layers to debond as the interface tensile strength is reached. For

the starting of interface debonding, normal stresses at the interface are divided in low tensile stresses at

midspan and high tensile stresses at the overlay’s edge, as shown in FIG.

Fig. 18 Evolution of the shear stress front along the interface for the REF specimens.

With the increase in load, the interface normal stresses vary in length between tensile and compressive,

according to discontinuities in the overlaid concrete layer due to cracking [ ]. These results are in

accordance with the vertical stress distribution between cracks shown in FIG. Debonding starts to occur at

the overlaid concrete edge, with the debonded portion identified by an almost null stress area (FIG).

For the specimens with reinforcement crossing the interface, stress distribution in the contact area was

changed due to the detailing of the interface. Specimens of the ANC detailing showed a concentration of

vertical tensile stresses in the anchored rebar region, with compressive stresses between this and the

midspan (FIG). For larger relative displacements, compressive stresses appear on one side of the

reinforcement as shown in FIG due to dowel action, which flexes the rebar.


49

Fig. 19 Vertical stresses at the interface for all detailings.

For the specimens of the STC detailing, stress distribution at the interface for larger relative

displacements also shows compressive stresses on one side of the steel connectors, also due to dowel

action, as illustrated on FIG. Since this reinforcement is distributed in rows at the interface, the change

between tensile and compressive stresses occur in each row of steel connectors. This behaviour can be

explained with the anchored reinforcement resisting the tensile strength required over the nominal

interface strength, as illustrated in FIG (retrieve schematics from third FCT report).

Fig. 20 Vertical stresses at the interface for all detailings before failure.

When analysing the cracking distribution inside the concrete layer, the principal compressive stress paths

can be identified in FIG. These paths show the concentration of stresses in the regions where

reinforcement in anchored, since the increase in tensile resistance of these strengthening elements

requires an equilibrium from compressive stresses.

Regarding the shear stress at the interface, an evolution is observed as the load increases, with a

maximum shear stress front evolving from the overlay’s edge to midspan, working with tensile stresses

and shearing the connection between layers. In FIG the evolution of the longitudinal shear stress τxz can

be observed, along with a stress direction change at the slab’s plastic hinge. This is coherent with the

schematics of FIG. (include in report, make new here…)

Fig. 21 Shear stresses at the interface for all detailings before failure.

Discussion Analysing the results, a good correlation is inherent to all FE models. This was possible due to calibration

of the software according to experimental data. The data mainly concerned material properties, but also

the relative displacements measured allowed for some tuning of the interface response. For interface


50

detailings with and without reinforcement, interface definition was the main parameter that affected

global behaviour and failure loads.

Stress distribution for the interface was numerically assessed in this work, with good correlation to

previous theoretical determination [ ]. Stress fields alternating between tensile and compressive at the

interface along with maximum shear stresses at the overlaid concrete edge could be observed. This is due

to the geometrical discontinuity for this region, which helps stresses to concentrate and shearing of the

interface.

The relationship between alternating stress fields and the position of the steel connectors did not verify

the theoretical assumption that the increased tensile resistance of the rebar crossing the interface would

lead to only tensile stresses near these elements. Although the stress distribution at the interface depends

on the detailing for the reinforcement crossing the interface, it also depends on the geometry of the slabs

and the loading conditions. This can be observed in FIG for the intermediate row of steel connectors,

where stresses are exclusively compressive near these elements.

Fig. 22 Compressive vertical stresses near the shear connectors.

Stresses for the longitudinal reinforcement could be analysed and compared with the experimental

measurements. A peak can be observed at the overlaid concrete edge for the non-linear stage of loading

for the STANC detailing. This behaviour could be explained through the larger rebar diameter anchoring

the overlaid concrete layer at the edge, when the contact interface had already deteriorated.

Dowel action of the reinforcement crossing the interface could be observed for the FE models. This

attests for the expected behaviour of the slabs where this particular resisting mechanism is accounted for.

Stress distribution for the steel connector resulted according to [Randl] and can be observed in FIG. The

maximum compressive stresses occurred until a depth “n”, with bending stresses changing the signal

below this depth.


51

Fig. 23 Normal stresses for the shear connectors sign changes at half the embedment length.

Cracking at the anchored longitudinal rebar zones resulted asymmetrical due to the bending nature of

loading, observed in FIG.

Fig. 24 Cracking around the anchorage of reinforcement crossing the interface.

When analysing the direction of the principal compressive stress inside the slabs… For the STANC

detailing of the interface, a clear concentration of stresses that lead to cracking occurred near the

reinforcement crossing the interface. A path is identified for these stresses where they cross the last row

of steel connectors up to the overlaid concrete rebar, and equilibrium occurs with the anchoring of these

elements. A strut and tie model can be defined for these slabs according to FIG.


52

Figure 13 - Principal compressive stress path and concrete cracking

The cracking pattern observed for this detailing of the interface was coherent with the inclined cracks

observed in the anchored rebar regions for the experimental specimens. This behaviour can be observed

in FIG, where an horizontal crack shifts from the interface to the rebar.

Figure 14 - Inclined cracks for the anchored rebar region (STANC detailing)

The lack of a clear and accountable intersection between both layers required the calculation of interface

stiffnesses regarding the adjacent materials fracture energy. For this method, a negative theoretical

displacement in the bond stress-slip relationship of the interface was considered, with the stiffness

calculated according to a linear stress growth until maximum stress was attained. This is the energy

needed for the first crack to appear, with the softening of stresses beginning right after, until degradation

of the interface and no stress is transferred. Although this method allowed for a good agreement with

experimental results, it is still an approximate way of calculating interface stiffness and cracking of the

interface.

The lack of an interfacial shear strength test did not allow experimental quantification of this parameter.

Mohr-Coulomb theory was then used to calculate a cohesion stress that allows the definition of the failure

surface characteristic for this theory, since it was a necessary parameter for the correct definition of the

interface. Good correlation was found with experimental results, but still just a calculated parameter with

no experimental assessment.

Interfacial debonding occurred for the same loads as the experimental tests, attesting for good correlation

with the experimental specimens. The softening of stresses at the interface after debonding defined

according to Reinhardt’s hyperbolic model [ ]. This resulted in a good correlation with experimental test

results, since the observed relationship between force and interface relative displacements was also

hyperbolic. Even when no reinforcement crossed the interface this evolution seemed to be characteristic

for and interface in the tensile face, according to FIG.


53

Structural failure was divided in two main modes, bending and shear, with the latter being predominant

due to the slender nature of the slabs with high amounts of reinforcement. Stress distribution due to

reinforcement crossing the interface allowed for a shear crack to develop before bending failure, which

led to brittle failure of the specimens. Only specimens with no reinforcement crossing the interface

resulted in bending failure through yielding of the substrate’s longitudinal rebar.

Conclusions The numerical FE models analysed allowed to verify the stress distribution for the interface and also the

stress fields inside the concrete layers. The latter allows for the identification of the main stress paths,

which load the interface and impact stress distribution. This was the main topic to be analysed since no

access and no measuring are provided at the interface.

Analysing all results from the numerical FE modelling, a good correlation with the experimental

specimens tested was attained, with the behaviour observed during these tests occurring in the

computational environment. Numerical FE models stress-deformation curves adjusted properly to

experimental specimens, with differences smaller than 10 %.

A key aspect for this type of structural strengthening was the interface between the two concrete layers,

which determined how the structure responds to the loads applied, and both local and global behaviour

were affected. Definition of this parameter determined how the slabs resisted the flexural and shear loads

imposed, mainly affecting its stiffness, which depends on the geometry and distribution of rebars.

Interface maximum stresses were set while defining the numerical model geometrical and material

parameters. This affected interface resisting strength, which behaved according to experimental specimen

testing.

A method for defining the interface normal and tangential stiffnesses was determined, which in turn

depends on the weakest concrete fracture strength. This defines the interface fracture strength since there

is no well defined line for the interface and no bonding agent. The weakest material rules the interface

strength, with the fracture energy affecting the strength of the entire structure. The same method was

applied to the grout-concrete interface for the reinforcement embedded in the concrete substrate with

good results.

Both debonding of the overlaid concrete layer and global failure, flexural and in shear, were found in the

numerical FE models with good correlation to the experimental tests. The debonding phenomenon

allowed to analyse the stress distribution at the interface, mainly in terms of the tensile and compressive

stress equilibrium. For the specimens with reinforcement crossing the interface, these elements affected

stress distribution, leading to greater tensile and consequently compressive stresses. Shear stress at the

interface also benefited from the increase in normal stresses because of the shear friction mechanism.

This allowed for a greater shear stress, and consequently larger relative displacements until debonding,

for an improved ductility of the interface.


54

6. Punching tests for bi-directional reinforced flat slabs strengthened with BCO

Experimental research

Strengthened test specimens definition

The experimental programme for assessing the performance of the strengthening method

presented in this work consisted of four reinforced concrete flat slabs reduced scale specimens,

strengthened with an overlaid reinforced concrete layer in the slab-column connection. These

were divided in a reference specimen with only surface preparation and no interface

reinforcement (S-REF), and the other three with different detailing for this reinforcement:

1. Steel connectors symmetrically distributed across the interface, embedded 80 mm with

cement grout in the existing substratum (S-STC);

2. Longitudinal reinforcement of the overlaid concrete anchored in the ends by embedding

80 mm in the existing substratum, with cement grout (S-ANC);

3. A combination of the two aforementioned detailings of the interface (S-STANC).

Due to referencing of the edge lifting phenomenon in several works [ , , ] and its inclusion in

design codes [MC2010, etc], detailing with reinforcement crossing the interface near the new

layer ends directly addresses this phenomenon, controlling the interface crack opening. Surface

preparation with an electric hammer and steel moil point was common for all strengthened

specimens since it has been proven to boost the performance of strengthening systems based on

new-to-old concrete interaction.

Dimensions for the test specimens were defined for a standard test setup used at the NOVA

University in Portugal, which tests the punching strength for a 5.00 m span flat slab. The reduced

scale specimen’s substratum comprised the area dimensions of 2300x2300 mm2, with 150 mm in

thickness. The strengthening intervention of overlaid concrete was calculated at 4d of the column

face, in order to comprise enough area for the punching shear crack to develop and anchoring of

the longitudinal reinforcement through bonding to surrounding concrete. An effective height d of

175 mm was calculated to account for intended punching load, and estimated tensile and shear

stresses at the concrete-to-concrete interface. For that, a 60 mm overlaid reinforced concrete


55

layer was determined with area dimensions of 1700x1700 mm2, which totals 210 mm thickness

of strengthened slab, with both layers illustrated in FIG.

Fig. 26 Detailing for the reference (S-REF) specimen, substratum (left) and overlay (right).

These dimensions would allow anchoring the longitudinal reinforcement for the S-ANC and S-

STANC detailings exactly at 4d of the column face, with a lateral concrete cover of 50 mm at the

layer’s ends. Longitudinal reinforcement calculated for the intended behaviour consisted in an

orthogonal mesh of 16 mm and 10 mm diameter rebars for the test specimen’s substratum top

and bottom layers. This yields a reinforcement ratio of 1.68 % for the substratum alone, which is

relatively high for current flat slab design, but still chosen due to exploratory work in 15 year old

flat slabs available with this reinforcement detailing. A minimum concrete cover of 20 mm was

determined for all rebar on both layers. For the overlaid concrete longitudinal reinforcement, an

orthogonal mesh of double 10 mm diameter rebars were calculated, considering a minimum

spacing to the interface and irrespective concrete cover. A resulting reinforcement ratio of 2.00

% for the strengthened cross-section presented a 20 % increase over the existing slab

reinforcement ratio.

Detailing of interface crossing reinforcement with steel connectors was determined for the

tensile stress that results from the punching load, as well as the shear stress from differential

stresses at longitudinal reinforcement. The two-dimensional character of flat slabs means that

some shear connectors will be loaded in the vertical, and both horizontal directions. (…, no FIG

reference yet)

Strain gauges

0d

2d

4d

#Ø16//0.100m

Strain gauges

0d

2d

#2Ø10//0.100m


56

Fig. 27 Directions of loading at steel connectors (S-STC detailing).

Anchoring of the longitudinal reinforcement in the overlaid concrete ends used the double rebar

detailing, with only one end of each rebar embedded in the substratum. A representation of all

detailing characteristics for reinforcement crossing the interface is illustrated in FIG for the S-

STANC specimen. The two other interface detailings (S-STC and S-ANC) are achieved by

eliminating the anchoring of the longitudinal reinforcement or the steel connectors, respectively.

Fig. 28 Detailing and strain gauge placing for the S-ANC (left) and S-STANC (right) test

specimens.

In FIGS the strain gauge placing can be identified. The main objectives were to capture the

greatest strain at the rebars and the differential strain between layers. The former was achieved

by placing the strain gauges in the column face, and the latter with coincidental placing in both

layers. Due to symmetry of the specimens, placing of the strain gauges were limited to two

quadrants.

Surface preparation of the substratum was performed with an electric chipping hammer and steel

moil point. Although its disadvantages regarding bruising and microcracking of the existing

Strain gauges

0d

2d

4d

#2Ø10//0.100m

Strain gauges

0d

2d

4d

#2Ø10//0.100m


57

concrete are well known, this technique is still the most widely used in strengthening or

retrofitting of concrete structures due to its low-level operation required, practical, and

economical characteristics. The resulting surface roughness was performed for each specimen

with concern that the moil tip should not go deeper than 10 mm at each stroke, thus protecting

the longitudinal reinforcement of the substratum and the minimum surrounding concrete cover.

The resulting surface can be observed in FIG for the reference specimen S-REF, which

represents the general aspect of surface preparation for all specimens. Surface assessment was

performed with a mechanical stylus type setup, with a displacement transducer TML CDP-50

connected to a HBM Spider 8 datalogger for data acquisition. Other methods for surface

assessment [REF] could not be used due to unavailability at the time or not recommended due to

the great roughness of the resulting surface, like the sand patch method [REF]. Results for

surface data regarding average and total roughness can be observed in TAB.

Table 4 - Average and total roughness parameters.

S-REF S-STC S-ANC S-STANC

Ra [mm] 3.8 3.3 3.1 3.6

Rt [mm] 9.2 7.4 8.3 9.1

Test setup

The punching test setup used, recurrent in NOVA University for this type of test and illustrated

in FIG, with test specimens monotonically loaded by a 1000 kN hydraulic jack ENERPAC RCH-

1003, positioned at the geometric center.

Fig. 29 General test setup, real life and schematics.

Loading was performed in the upwards direction by a steel plate with 200x200 mm2 area, and 50

mm thickness. Eight steel plates with 100x100 mm2 in area and 20 mm thickness guaranteed


58

kinematic boundary conditions at the zero moment line of the specimens. These elements were

linked to spreader beams with 15.2 mm high strength strands, and the whole support setup was

fixed to the laboratory strong floor by high strength 40 mm threadbars. All elements are

illustrated in FIG.

Fig. 30 General dimensions for the test setup, top view (left) and side view (right).

Punching load was measured at the supports by eight TML CNC-200KNA load cells. Vertical

displacements were measured by fourteen 100 mm TML CDP-100 displacement transducers,

thirteen on the top face and one on the bottom face, and four 50 mm TML CDP-50 displacement

transducers at the supports, illustrated in FIG.

Fig. 31 Vertical and relative displacement measuring instrumentation layout.

D7

D6

D5

D1

R3

D11 D12 D13 R2

rsub(dsub)

rolay

(dolay)

Zero Moment

line

D14 D18

D10 D2 D11 D12 D13

R2

R1

CC2 CC3


59

Relative displacement between layers was also measured horizontally with four 50 mm TML

CDP-50 displacement transducers (R series). Data acquisition was performed with Spider 8 and

Centipede dataloggers, both from HBM, all monitored by HBM Catman V6.0 software. Loading

until failure of the specimens was electronically controlled by a WALTER+BAI PKNS19D

hydraulic pump, at a speed of 0.25 kN/s.

Materials characterization

Material characterization was performed for different rebar sizes, grout for anchoring the rebar

crossing the interface, concrete compressive and tensile strength for both layers, and pulloff

strength at the interface between layers. Rebar testing was performed with a INSTRON 8874

tensile test machine, according to EN 10002-1 guidelines, with test results for yield and tensile

strength in TAB. The mechanical properties of the cement grout were characterized according to

RILEM PC-5 and PCM-8 guidelines for compressive and tensile strength, respectively. Bond

strength was also assessed with pullout tests for the irrespective rebar sizes. Material parameters

for the cement grout were a flexural tensile strength of 9.7 MPa, a compressive strength of 78.8

MPa, and bond strength of 16.2 MPa. The concrete characteristics were evaluated according to

EN 12390-3 for the compressive strength, and EN 12390-6 for the splitting tensile strength. The

former was carried out on 150 mm cubic specimens, and the latter on 150x300 mm cylinders.

Pulloff strength was carried out with 150x150 mm2, where the bottom layer portion of the

specimen was isolated 10 mm in depth, and the top layer, 60 mm thick, was bonded during

casting of the overlaid concrete. Pulloff tensile strength results can be observed in TAB.

Table 5 – Strength characteristics for the steel bars.

Ø6 mm Ø10 mm Ø16 mm

fy [MPa] 541.0 530.6 ###.#

fu [MPa] 692.7 627.5 ###.#

fy – mean yield stress of steel.

fu – mean tensile strength of steel.

Table 6 – Concrete and bond strength characteristics.

Reference

(S-REF)

Steel

connectors

(S-STC)

Anchored

rebar

(S-ANC)

Steel connectors

+ Anchored

rebar

(S-STANC)

Substrate fc,cube [MPa] 32.8 26.4 34.8 25.6

fc,t [MPa] 2.9 2.6 3.0 2.3


60

Bond σpull-off [MPa] 0.85 1.04 0.82 0.80

Overlay fc,cube [MPa] 36.9 34.3 37.2 39.3

fc,t [MPa] 2.9 3.2 2.6 2.8

σpull-off – mean value for the pull-off tensile strength of the interface between layers.

fc,cube – mean value for the compressive strength of concrete in cubic specimens.

fc,t – mean value for the tensile strength of concrete.

Experimental results

Failure modes

The failure modes were identified for all specimens, with punching shear failure for all

specimens with reinforcement crossing the interface, and debonding of the top layer for the

reference specimen. Failure for the latter occurred by debonding of the top layer at the interface,

as illustrated in FIG.

Fig. 32 Failure and crack pattern for the S-REF specimen.

During testing of this specimen, radial cracks could be observed on the top face of the overlaid

concrete, which could indicate a significant amount of stresses were being transferred to the top

layer. When no more deformation could be supported by the interface, debonding occurred.

These cracks never evolved from hairline cracks, and closed right after debonding of the overlaid

concrete layer. Analysing the resulting surface after debonding, it can be concluded that the top

layer debonded clean from the bottom layer, with the original surface after preparation with

almost no visible alterations. Removing the overlaid concrete layer, the punching shear cone that

developed in the substratum can be observed, as illustrated in FIG.


61

Fig. 33 Interface surface for both layers after debonding.

Where some chunks of concrete from the substratum can be found attached in the overlaid

concrete surface, ridges remained intact for this layer, which is consistent with the concrete

strength assessed for each layer. The detailing with shear connectors allowed for the punching

shear failure to reach the top face of the overlaid concrete, as illustrated in FIG. It should be

noted that the cracking pattern due to punching never reached the edges of this layer, and dowel

action occurred for the longitudinal reinforcement.

Fig. 34 Failure and crack pattern for the S-STC detailing.

The most prominent tangential crack can be observed at two times the substratum effective

height (≈0.25 m), and visually the damage reached a distance that was 5 cm from the overlaid

concrete layer edges. Failure mode was similar for the detailing with the longitudinal

reinforcement anchored in the substratum, as observed in FIG.

≈0.25m

≈0.25m


62

Fig. 35 Failure and crack pattern for the S-ANC detailing.

For the specimen with all solutions combined the failure mode followed that of the latter ones,

with punching failure reaching the top face of the overlaid concrete layer, as observed in FIG.

Similar to those specimens, main tangential cracking was also at 0.25 m from the column face

(FIG).

Fig. 36 Crack pattern for the S-STANC detailing.

Debonding and failure loads

Loading until failure was performed in a monotonic manner for all specimens, with debonding of

the overlaid concrete layer for the reference specimen, and punching failure for the specimens

with reinforcement crossing the interface. Failure for the latter specimens was characterized by a

residual resisting capacity for larger displacements about two times that of the reference

specimen. The relationship between load and deflection at the center of each specimen can be

observed in FIG.

≈0.25m

≈0.25m ≈0.25m


63

Fig. 37 Load deflection relationship at the center of the test specimens.

The load-deflection relationships presented above show a more ductile transition from punching

load to residual load when compared to the reference specimen. Analysing the failure loads for

all specimens one can verify the registered values are of the same magnitude. Considering the

predicted loads provided by the Model Code 2010 for the substratum alone, the load increment

due to strengthening with a concrete overlay was greatest for the specimens with shear

connectors, as observed in FIG. A 65 % increment for the S-STC detailing and 62 % increment

for the S-STANC detailing were the largest registered values. This phenomenon attests the good

performance for the detailing of the shear connectors stitching the two layers.

0

200

400

600

0 10 20 30 40 50

Lo

ad [

kN

]

Deflection [mm]

Load-deflection REF slab

0

200

400

600

0 10 20 30 40 50

Lo

ad [

kN

]

Deflection [mm]

Load-deflection STC slab

0

200

400

600

0 10 20 30 40 50

Lo

ad [

kN

]

Deflection [mm]

Load-deflection ANC slab

0

200

400

600

0 10 20 30 40 50

Lo

ad [

kN

]

Deflection [mm]

Load-deflection STANC slab


64

Fig. 38 Experimental failure loads divided by layer for all specimens.

Analysing the load-deflection relationship for each of the quadrants of the slab specimens, one

can evaluate the measured perimeters and attest the distance of the critical shear crack from the

column face. The difference in stiffness becomes apparent between the first (≈0.225 m) and the

second (≈0.550 m) measured perimeters, with the distance of two times de substratum effective

height (≈0.250 m) in-between. (maybe FIG).

Cracking is not apparent in the load-deflection relationships in FIG, and a progressive decrease

in stiffness can be observed until sudden failure by punching of the specimens. Cracking is only

apparent in the instrumented rebar, with a sudden increase in steel strain for a load about 100 kN

for all specimens, observed in FIG. This increase can be explained by the sudden transfer of

tensile stresses from concrete to the reinforcement upon cracking.

0

100

200

300

400

500

600

1 2 3 4

Lo

ad [

kN

]

Experimental failure loads

Overlay Substratum

S-ANC S-STANC

f c=

26

,2 M

Pa

f c

=29,5

MP

a

559,8 kN

f c=

27

,9 M

Pa

f c

=29,7

MP

a

567,6 kN

f c=

21

,1 M

Pa

f c

=27,5

MP

a

535,8 kN

f c=

20

,5 M

Pa

f c

=31,5

MP

a

549,8 kN

S-REF S-STC

0

200

400

600

0 500 1000 1500 2000

Lo

ad [

kN

]

Strain [µm/m]

REF - higher effective height

Substratum

Overlay0

200

400

600

0 500 1000 1500 2000

Lo

ad [

kN

]

Strain [µm/m]

STC - higher effective height

Substratum

Overlay


65

Fig. 39 Load deflection relationship at the center of the test specimens.

The latter strain was measured in the column face and states the maximum strain reached for all

specimens. Further away from the column face, the increase in steel strain depends on cracking

of the specimens, not following a trend in load values, but can happen as late as 300 kN.

Yielding of the reinforcement was not reached for any specimen, with a maximum value

registered of 1.89 ‰ for the S-STC specimen. One can notice that a strain about 1.50 ‰ was

registered at the maximum load for the S-REF, S-ANC, and S-STANC specimens. A

phenomenon is also identified for all specimens, consisting of a logarithmic variation of strain at

two times the strengthened cross section effective height (≈0.350 m).

7. Thesis and Dissertations

Università Degli Studi Firenze From the contribution of the foreign students in the experimental campaign for the bi-directional flat

slabs, a dissertation was produced with the results, and presented at the institution on November 25th, for

the attainment of the master degree for the three students. The thorough analysis that was performed

based on the test results allowed for a structured presentation and interpretation of the results, that will be

jointly conducted to a scientific communication to a recognized international journal.

0

200

400

600

0 500 1000 1500 2000

Lo

ad [

kN

]

Strain [µm/m]

ANC - higher effective height

Substratum

Overlay0

200

400

600

0 500 1000 1500 2000

Lo

ad [

kN

]

Strain [µm/m]

STANC - higher effective height

Substratum

Overlay


66

Fig. 40 Presentation for the attainment of the Master’s Degree at U. Firenze.

8. Communications in conferences

SILE2015

A presentation was submitted and carried out by professor Válter Lúcio on this year’s “Seminário

Internacional de Ligações Estruturais”. The main focus of this presentation was to analyse the behaviour

and structural strengthening performance of the BCO technique. It also presented some work that was

submitted to the “Engineering Structures” journal, published by Elsevier.


67

Fig. 41 Presentation at the SILE 2015 seminar.

fib – Cape Town 2016

An oral communication was accepted to the fib Symposium to be held in Cape Town later in 2016, with

the following abstract:

“Strengthening of concrete structures with a new concrete layer applied on the tensile face has been

proved effective for service and ultimate limit states. Although it requires the consideration for the

premature debonding, correctly designed it stands as an efficient and economical method for structural

strengthening.

Different detailing solutions for the rebars crossing the interface between the two concrete layers were

tested and analyzed in order to identify the main factors that influence the behavior of the composite

section. The knowledge from this study was applied to the strengthening of flat slabs, where large stress

concentration occurs. The usual brittle failure of these zones by punching further influence the behavior

of the interface between the two concrete layers.

This paper presents the study performed on concrete unidirectional slabs and on flat slabs strengthened

with a bonded concrete overlay. The experimental research consisted on twelve specimens loaded

monotonically and concentrically, which were then compared to numerical models analyzed with ATENA


68

3D® software. All results were then compared with current codes and regulations, namely the Eurocode

2 and the Model Code 2010.”


69

9. References

[ 1 ] A. Pinho Ramos, Válter J.G. Lúcio, Duarte M.V. Faria, The effect of the vertical

component of prestress forces on the punching strength of flat slabs, Engineering

Structures, Volume 76, 1 October 2014, Pages 90-98, ISSN 0141-0296.

[ 2 ] Clément, T.; Ramos, A.; Fernández Ruiz, M. and Muttoni, A. – Design for

punching of prestressed concrete slabs. Structural Concrete, 14: 157–167. doi:

10.1002/suco.201200028, 2013.

[ 3 ]

Faria, D.; Lúcio, V. and Ramos, A. – Post-Punching Behaviour of Flat Slabs

Strengthened with a New Technique using Post‑Tensioning, Engineering

Structures, Volume 40, pp. 382-397, July, 2012.

[ 4 ]

Faria, D.; Biscaia, H.; Lúcio, V. and Ramos, A. – Material and geometrical

parameters affecting punching of reinforced concrete flat slabs with orthogonal

reinforcement. Short Paper, fib Symposium PRAGUE 2011 – Concrete

Engineering for Excellence and Efficiency, Prague, June, 2011.

[ 5 ]

Faria, D.; Biscaia, H.; Lúcio, V. and Ramos, A. – Punching of reinforced concrete

slabs and experimental analysis and comparison with codes. Proceedings of

IABSE-Fib Codes in Structural Engineering – Developments and Needs for

International Practice, Cavtat, Dubrovnik, May, 2010.

[ 6 ]

Faria, D.; Inácio, M.; Lúcio, V. and Ramos, A. – Punching of Strengthened

Concrete Slabs – Experimental Analysis and Comparison with Codes, IABSE,

Structural Engineering International, No. 2 – “Codes of Practice in Structural

Engineering”, May, 2012.

[ 7 ] Faria, D.; Lúcio, V. and Ramos, A. – Reforço de lajes com recurso a pós tensão

com ancoragens por aderência, Encontro Nacional Betão Estrutural 2012,

Faculdade de Engenharia da Universidade do Porto, October,2012.

[ 8 ] Faria, D.; Lúcio, V. and Ramos, A. – Development of a Design Proposal for a

Slab Strengthening System using Prestress with Anchorages by Bonding,

Proceedings of fib symposium Tel Aviv 2013, Tel Aviv, April, 2013.

[ 9 ] Faria, D.; Lúcio, V. and Ramos, A. – Pull-out and push-in tests of bonded steel

strands. Magazine of Concrete Research, Thomas Telford, Volume 63, Issue 9,

pp. 689-705, September, 2011.

[ 10 ] Faria, D.; Lúcio, V. and Ramos, A. – Strengthening of flat slabs with post-

tensioning using anchorages by bonding. Engineering and Structures, Volume

33, pp. 2025-2043, June, 2011.

[ 11 ] Faria, D.; Lúcio, V. and Ramos, A. – Strengthening of Reinforced Concrete Slabs

Using Post Tensioning with Anchorages by Bonding. Proceedings of fib

Symposium Concrete:21st Century Superhero, London, June, 2009.


70

[ 12 ]

Faria, D.; Lúcio, V. and Ramos, A. – Bond Behaviour of Prestress Steel Strands

Bonded with an Epoxy Adhesive and a Cement Grout for Flat Slab Strengthening

Purposes – Experimental Study. Proceedings of 3rd fib International Congress,

Washington, May/June, 2010.

[ 13 ] Gomes, J. and Ramos, A. – Estudo Experimental do Punçoamento em Lajes

Reforçadas com Armadura Transversal Aderente Pós-Instalada, Encontro

Nacional Betão Estrutural 2010, Lisbon, November, 2010.

[ 14 ] Gomes, J. and Ramos, A. – Punçoamento em Lajes Fungiformes Reforçadas

com Parafusos Transversais Aderentes (Parte 1). Revista Internacional

Construlink, Nº 30, Vol. 10, 23-33, June, 2012.

[ 15 ] Gomes, J. and Ramos, A. – Punçoamento em Lajes Fungiformes Reforçadas

com Parafusos Transversais Aderentes (Parte 2). Revista Internacional

Construlink, Nº 30, Vol. 10, pp. 34-43, June, 2012.

[ 16 ] Gouveia, N.; Fernandes, N.; Faria, D.; Ramos, A. and Lúcio, V. – Punching of

Steel Fibre Reinforcement Concrete Flat Slabs, Proceedings of fib symposium

Tel Aviv 2013, Tel Aviv, April, 2013.

[ 17 ]

Inácio, M.; Ramos, A. and Faria, D. – Strengthening of flat slabs with transverse

reinforcement by introduction of steel bolts using different anchorage

approaches. Engineering and Structures, Volume 44, pp. 63-77, November,

2012.

[ 18 ]

Inácio, M.; Ramos, A.; Lúcio, V. and Faria, D. – Punçoamento de lajes

fungiformes reforçadas com parafusos – efeito da área e posicionamento da

ancoragem, Encontro Nacional Betão Estrutural 2012, Faculdade de Engenharia

da Universidade do Porto, October, 2012.

[ 19 ] Inácio, M.; Ramos, A.; Lúcio, V. and Faria, D. – Punching of High Strength

Concrete Flat Slabs - Experimental Investigation, Proceedings of fib symposium

Tel Aviv 2013, 4p, Tel Aviv, April, 2013.

[ 20 ]

Mamede, N.; Ramos, A. and Faria, D. – Experimental and parametric 3D

nonlinear finite element analysis on punching of flat slabs with orthogonal

reinforcement. Engineering and Structures, Volume 48, pp. 442-457, March,

2013.

[ 21 ]

Mamede, N.; Ramos, A. and Faria, D. – Análise do efeito de características

mecânicas e geométricas que afetam o comportamento ao Punçoamento de

lajes fungiformes, Encontro Nacional Betão Estrutural 2012, Faculdade de

Engenharia da Universidade do Porto, October, 2012.

[ 22 ]

Micael M.G. Inácio, André F.O. Almeida, Duarte M.V. Faria, Válter J.G. Lúcio,

António Pinho Ramos, Punching of high strength concrete flat slabs without

shear reinforcement, Engineering Structures, Volume 103, 15 November 2015,

Pages 275-284, ISSN 0141-0296.


71

[ 23 ]

Nuno D. Gouveia, Nelson A.G. Fernandes, Duarte M.V. Faria, António M.P.

Ramos, Válter J.G. Lúcio, SFRC flat slabs punching behaviour – Experimental

research, Composites Part B: Engineering, Volume 63, July 2014, Pages 161-

171, ISSN 1359-8368.

[ 24 ] Paias, J. and Ramos, A. – Estudo Experimental do Punçoamento em Lajes de

Betão Reforçado com Fibras de Aço, Encontro Nacional Betão Estrutural 2010,

Lisboa, November, 2010.

[ 25 ] Ramos, A. and Lúcio, V. – Post-Punching Behaviour of Prestressed Concrete

Flat Slabs. Magazine of Concrete Research, Thomas Telford, 60, no. 4, May,

2008.

[ 26 ] Ramos, A. – Punçoamento em Lajes Fungiformes Pré-Esforçadas. Tese

apresentada no Instituto Superior Técnico, Universidade Técnica de Lisboa para

obtenção do Grau de Doutor em Engenharia Civil, March, 2003.

[ 27 ] Ramos, A., Lúcio, V., Faria, D. e Inácio, M. – Punching Research at

Universidade Nova de Lisboa. Design Of Concrete Structures and Bridges Using

Eurocodes, Bratislava, September, 2011.

[ 28 ] Ramos, A.; Lúcio, V. and Regan, P. – Punching of flat slabs with in-plane forces,

Engineering Structures, Volume 33, Issue 3, March, 2011.

[ 29 ] Silva, R.; Faria, D.; Ramos, A. and Inácio, M. – A physical approach for

considering the anchorage head size influence in the punching capacity of slabs

strengthened with vertical steel bolts, Structural Concrete, June, 2013.

reforço de lajes fungiformes com recurso a …...with hydro demolition with high-pressure water...

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