bal canti

CONNECTION OF BALANCED CANTILEVER BRIDGES WITH NEIGHBOR-HOODING TUNNELS

Stergios A. Mitoulis, Ioannis A. Tegos Aristotle University, Department of Civil Engineering, Thessaloniki, Greece,

[email protected]

ABSTRACT

It is known that a large number of bridges are constructed between tunnels, whose ends are close to the end parts of the bridge deck. This co-existence can be devel-oped in order to reduce the structural cost of bridges, as the end parts of these bridges can be restrained by the tunnels. By extension, a complex system, which consists of three discrete structures, can be developed in order to enhance the seis-mic performance of the bridge. However, the restrain of the end parts of long bridges induces serviceability problems, which are possible to be accommodated by means of appropriate approach elements and expansion joints. The aforemen-tioned approach slabs are connecting the bridge with the tunnels. In the present study a new type of a connective approach element is proposed and is parametri-cally investigated. The proposed continuity slab is designed in order to accommo-date both serviceability and earthquake resistance of the bridge. The attempt does not only serves cost-effectiveness but also esthetics, as the reduction of the cross sections of the piers is related to slim and attractive side views of bridge structures.

KEYWORDS

Bridge, seismic, serviceability, balanced cantilever, tunnel, complex, interlocking.

1 INTRODUCTION

The construction of bridges with balanced cantilevers is a popular structural method, as this method has great advantages over other structural methods, temporary shoring would disrupt traffic and service below and falsework would not only be expensive but also a hazard [1]. On the other hand, balanced cantilever bridges improve es-thetics, as their lines provide, through their continuity, a visual connection between related bridge parts. The central span of prestressed concrete cantilever bridges can

472 S.A. Mitoulis, I.A. Tegos

extend almost 300 m as the longest central span of Stolmasundet bridge, which links the islands of Stolmen and Sjelborn in Austevoll Norway is 301 m long. Most of the past and current studies on balanced cantilever bridge construction refer to the constructability and to the resistance of these bridge systems against static combinations, as well as to the redistribution of actions due to creep, shrink-age and relaxation [1-4]. However, in areas with high seismicity, balanced cantile-ver bridges usually respond with large seismic displacements due to the flexibility of their resisting system. The “key point” for the enhancement of large bridge dis-placements is usually the increase in the damping, which usually leads to the use of viscous dampers and high damping bearings [5]. Also the development of the hys-teretic behavior of the piers, since in most cantilever bridges the piers are con-nected to the deck in a monolithical way, can lead to a significant decrease in the seismic actions [6] and by extension, in the reduction in the seismic displacements. The present study proposes an unconventional seismic design technique for canti-lever bridges, in the sense that the seismic actions are not only enhanced by the conventional seismic resisting elements – piers and seismic isolation devices – but also by external elements which participate during earthquake. The proposal is based on the observation that the geomorphology, which leads to the construction of cantilever bridges, often requires the construction of tunnels, which are neighbor-hooding on both sides of the cantilever bridge. The co-existence of these transporting systems can be developed in order to utilize the stiffness of the tunnel and to reduce the horizontal seismic movements – longitudinal and transverse – of the bridge, by “connecting” the bridge with the tunnels. The proposed “connection” can be implemented in all reinforced concrete bridges, which are lying between tunnels. It is noted that the monolithical connection is not always possible, as the great total length of the bridge in combination with creep, shrinkage [7] and ther-mal effects [8] usually lead to large displacements of the end parts of long bridges, and by extension the continuity slab, which would connect the deck with the tunnel would probably fail in tension or compression in-service. The present investigation provides a semi-connection, as appropriate gaps between the connecting slab and the foundation of the tunnel provide the appropriate space for the deck to develop its constraint movements.

2 DESCRIPTION OF THE “REFERENCE” BRIDGE

The present study used a bridge of Egnatia Odos Motorway [9], which was consid-ered to be the “reference” bridge (Fig. 1). This conventional bridge was built by the balanced cantilever method, has three spans and a total length equal to L=349.0 m. The first and the third span have a length equal to 94.0 m while the central and longest span is 160.0 m long. The bridge is located at Malakasi-Grevena area of Egnatia Odos Motorway. The deck of the bridge consists of a continuous box-girder, which is monolithically connected to the two piers of the bridge, while is seated on two sliding bearings on the abutments. The height of the box section of the deck decreases from

Connection of balanced cantilever bridges with neighbor-hooding tunnels 473

9.00 m, (pier), to 3.50 m, (mid-span). The width of the deck is 13.60 m. The piers are rectangular hollow sections and the thickness of their webs is 0.80 m. The di-mensions of the piers are 6.00 m and 7.20 m in the longitudinal and in the transverse direction of the bridge correspondingly and the thickness of their web is 0.80 m. The first pier’s (P1) height is 58 m while the second one is 55 m. The piers are founded on circular wells which are 18.0 m deep with a diameter of 10.0 m. The abutment is a conventional seat type abutment, which provides the appropriate expansion joint [5] between the deck and its backwall, while it restrains the movement of the deck in the transverse direction as capacity design stoppers are installed on the abutments. The bridge is founded on ground type B [10] and the peak ground ac-celeration adopted was equal to ag = 0.16 g. The importance factor adopted was equal to γΙ = 1.30, while the behaviour factors were equal to qx = qy = 3.5 for the longitudinal, the transverse and the vertical direction of the bridge respectively.

3 DESCRIPTION OF THE PROPOSED INTERLOCKING SYSTEM

The proposed connecting slab has the ability to accommodate part of the induced constraint movement of the bridge’s deck, due to creep, shrinkage and prestressing. On the other hand, this semi-connection ensures the seismic participation of the rigid foundation of the tunnel and restrains the seismic movement of the deck in the longitudinal and in the transverse direction. The linking-key is connecting the con-tinuity slab with the foundation of the tunnel. The length of the continuity slab was considered to be equal to the common distance of the bridge abutment and the tun-nel, which is usually equal to 25.0 m. The dimensions and the reinforcement of the continuity slab should be appropriately selected in order to avoid the in-service buckling of this element. The undesirable in-service friction between this slab and the approach embankment is reduced as a slide-on slab is provided between the connecting slab and the backfill. The linking-key is formulated, on the one hand, by increasing the thickness of the connecting slab, and on the other hand, by an appropriate reformation of the tun-

A1

P1P2

A258.0m 55.0m

18.0m18.0m

94.5m 160.00 94.5m349.0

Fig. 1: Longitudinal section of the “reference” conventional bridge located in Malakasi-

Grevena territory of Egnatia Odos motorway.


nel’s foundation (Fig.2). Finally, the linking-key and the connecting slab are rein-forced with transverse reinforcement in order to withstand the high shear action, which is developed during earthquake. In order to formulate the needed embed-ment of the linking-key, an extension of the tunnel’s foundation towards the back-fill is needed. Also, the configuration of a U-shaped reception, in which the strengthened linking-key of the connecting slab is restrained, is needed (Fig. 2).

3.1 Determination of the widths of the expansion joints The system of the linking-key and the structural embedment provides two expan-sion joints, whose widths, ∆1 and ∆2, are accommodating only the in-service movement of the deck. It is noted that the conventional bridge design requires ex-pansion joints whose width is usually determined according to codes [5] and takes into account the in-service [11] and part of the seismic displacement actions. How-ever, the cost of providing a road joint to accommodate large seismic deflections may be prohibitive and a compromise is usually adopted [12]. Also, the minimiza-tion of the expansion joints is related to the objective of the present investigation that is the maximization of the seismic participation of the proposed restraining system, as smaller widths of the expansion joints lead to a more efficient reduction of the displacements of the deck [13] and by extension to the reduction of the struc-tural cost of the piers [14]. The main objective of the study was to define the opti-mum compromise between serviceability and earthquake resistance. In the present investigation the development of creep and shrinkage effects of the deck were taken into account by considering equivalent thermal variations according to PCI [7]. Also, a probabilistic approach of the value of the thermal movement [15] was enterprised in order to minimize the required widths of the expansion joints, ∆1 and ∆2.

stirrups

tunnel

shear key (bx by = 1000cm 250cm)embedment(Bx By = (1000+∆1+∆2cm) 250cm

tunnel foundation

continuity slabtslab=40cmsiding joint

deck

backwall

expansion joint

approachembankment

slide-on-slab

L key

∆1∆2shear key

embedmentdeck

abutment

Fig. 2: The proposed interlocking system, with the continuity slab and the linking-key.


According to the aforementioned determination of the widths of the joints, the deck of the bridge is not being compressed during its expansion, as joint ∆1 is not closed during the maximum expansion of the deck [15]. Joint ∆1 remains open even in the first years of bridge service, and while the creep and shrinkage, which result in the permanent contraction of the deck, have not been developed yet. More specifically, the width of joint ∆1, after the completion of the construction, can absorb the maxi-mum constraint movement of the deck due to its total thermal expansion [15]. After the first years of the service of the bridge, and while creep and shrinkage of the deck have almost been completed, joint, ∆1 remains open. The width of the second joint, ∆2, after the completion of the construction was adjusted by taking into ac-count the maximum contraction of the deck due to creep, shrinkage, prestressing and thermal contraction. The determination of ∆2 is related to three discrete design criteria: (a) the control of the maximum allowable tension of the deck slab, (b) the in-service allowable cracking of the connecting slab and, (c) the minimization of the width of the joint.

4 MODELING OF THE ANALYSED BRIDGE SYSTEMS

4.1 Modelling of the “reference” conventional bridge The seismic analysis of the bridge allows the use of simplified stick models, as the parameters, which influence the results of the analysis, are the stiffnesses the masses and the damping. In Figure 3 the stick model of the “reference” bridge is given. The deck of the bridge was modelled by 160 frame elements, which are taking into account the variation of the geometry of the deck’s cross section from the supports to the spans. The deck is supported on slide supports on the abutments, as the deck of the “reference” bridge is supported on two sliding bearings. Its trans-verse displacements were restrained, due to the existence of the transverse stop-pers. The piers and the well foundations were also modelled by frame elements. The resistance of the soil was taken into account by assigning linear spring ele-ments in the two horizontal directions and in the vertical direction. The stiffness values of these springs resulted from the geotechnical in-situ tests of the “refer-

deck (83 frame elements)

foundation well(18 frame elements)

hinges

(16 frame elements)pier Ρ2(16 frame elements)

pier Ρ1

soil springs

stiff zone

hinges

stiff zones4,60m

4,60mdeck

Fig. 3: The model of the “reference” conventional bridge.


ence” bridge. The possible plastic hinges that could be developed on the head and on the feet of the piers were determined by means of RCCOLLA [16].

4.2 Modelling of the proposed interlocking system The proposed interlocking system, which connects the bridge with the neighbor-hooding tunnels, consists of two discrete parts: (a) the continuity slab, which is the extension of the deck slab of the bridge onto the backfill and, (b) the linking-key, which consists of a stopper and its embedment in the foundation of the tunnel (Fig. 2). The continuity slab was modelled by a frame element and a multi-linear link in series. The cross section of the continuity slab has a thickness equal to 0.30 m and a transverse width equal to 8.0 m. The multi-linear link in series models the in-service axial response of the continuity slab, which acts as a tension-tie during the contraction of the deck, whereas it acts as a compression-strut during the expansion of the deck (Fig. 4). The axial resistance of the continuity slab during its tension was calculated by the axial stiffness of its longitudinal reinforcements, which on one hand ensure the in-service distribution of the cracks of the slab and on the other hand control the axial tension of the deck slab. The axial resistance of the continuity slab during its compression was determined by the axial stiffness of a frame element which has the cross section of the slab and a length equal to 25 m. It is noted that the continuity slab is compressed only during the seismic event, as the joint ∆1 ensures that the linking-key is not activated during the expansion of the deck of the bridge. The linking-key provides two joints whose widths, ∆1 and ∆2, (see Fig. 4) ac-commodate the expansion of the deck and part of the total contraction of the deck correspondingly. The linking-key was modelled by a multi-linear link which takes into account both of the gaps ∆1 and ∆2, Fig. 4(b) and Fig. 4(c). The afore-mentioned multi-linear link has three discrete branches: (i) a compression branch which models the case that the deck is moving towards the tunnel during earth-quake (the case of the compression of the deck during an extreme expansion of the deck was avoided through the appropriate selection of ∆1) and ∆1 is closed, (ii) a tension branch, which models the case that the deck is either drawing away from the tunnel during earthquake or is contracting during service, and ∆2 is closed and , (iii) a constant branch, which models the case that both gaps, ∆1 and ∆2, are open and by extension the bridge and the tunnel are not interacting. The stiffness of the compression and the tension branch correspond to the stiffness of the contact element, which models the collision between the tunnel and the conti-nuity slab. The stiffness of this contact element was determined by the axial stiff-ness of the slab of the deck [17, [18]. The widths of the joints, provided in the interlocking system, can vary, at the beginning of the seismic event, due to the expansion or the contraction of the deck and due to the permanent contraction of the deck, which is caused by creep and shrinkage effects. The present study con-sidered two different states of the joints at the beginning of the seismic event,


which correspond to the two extreme cases: (a) the maximum contraction of the deck, which corresponds to ∆1 = 0 and ∆2 = 11 cm and, (b) the maximum expan-sion of the deck, which corresponds to ∆1 = 5.2 cm and ∆2 = 5.8 cm. It is noted that the initial widths of these joints and while creep and shrinkage effects have not been developed are: (a) ∆1 = 4.4 cm, and covers the total expansion of the deck (b) ∆2 =6.6 cm and accommodates part of the deck’s constraint contraction due to creep, shrinkage, prestress and thermal contraction. The continuity slab accommodates the rest of the constraint contraction of the deck, as it can absorb a total of 3.0 cm due to its in-service allowable cracking.

4.4 Parameters of the study The proposed semi-connections of the bridge with the tunnels was parametrically investigated in order to identify the earthquake resistance efficiency of the system in terms of percentage reductions in deck’s displacements and in piers’ actions. The parameters of the study mainly resulted from the seismic action considered, namely the soil category and the peak ground acceleration. Specifically, the “refer-ence” and the modified bridge systems were subjected to artificial accelerogramms that were compatible to soil A, B and C dependent Eurocode 8 elastic spectra, [6], and two different peak ground accelerations, ag=0.16g and ag=0.24g were consid-ered. The non-linear response of the “reference” and the unconventional bridge was analyzed using the FEM code SAP 2000 ver. 11.0.0, [19]. Dynamic non-linear time history analysis was implemented and the direct integration of β-Newmark method was chosen.

u

Pstop.

∆1=11cm

∆2 closed

Maximum Contraction

u

Pstop.

∆1=5,2cm

Maximum Expansion

∆2=5,8cm

Resistance of the continuity slab (Lslab=25m)

axial deflectiony

Pslab

4 800KN

48 000KN

0.015mtension

compression

deck (83 frame elements)

foundation well(18 frame elements)

hinges

(16 frame elements)pier Ρ2(16 frame elements)

pier Ρ1

soil springs

stiff zone

hinges

Fig. 4: The model of: (a) the unconventional bridge, (b) the linking-key resistance for

the maximum contraction and (c) for the maximum expansion of the deck and (d) the continuity slab.

(a)

(b) (c)

(d)


5 RESULTS

The present study focuses on determining the earthquake resistance of an uncon-ventional interlocking system, which has the ability, on the one hand, to restrain the seismic movements of cantilever bridges, which are lying between tunnels (Fig. 1), and on the other hand to accommodate the serviceability of the deck of long bridges. The efficiency of the proposed system was assessed by calculating the percentage reductions in the longitudinal and in the transverse seismic movements of the deck. These reductions resulted from the strong participation of the tunnels’ stiff foundations, which was accomplished through the extension of the deck slab onto the backfills and the semi-connection of this slab with the tunnels by means of an appropriately configured linking-key. The percentage reductions in the deck’s seismic movements, which are expressed by Eq. 1, resulted from the comparison of the response of the “reference” and the unconventional bridge system.

, .

. . 1= − E,UNCONV.

E CONV

uP R

u (1)

where P.R. is the percentage reduction in the movements of the deck, (lon-gitudinal or transverse),

uE,UNCONV. is the seismic displacement of the deck of the unconventional bridge and

uE,CONV. the seismic displacement of the deck of the conventional bridge.

From Eq. 1 it can be extracted that if P.R.>0 then the unconventional bridge system responds with smaller displacements and by extension the proposed restraining system is efficient, while if P.R.<0 then the proposed system is considered to be insufficient. Figure 5 shows the variation of P.R. factor for the case of a Seismic Zone I, that corresponds to peak ground acceleration equal to ag = 0.16 g. The horizontal axis corresponds to the deck joints above its sequential supports i.e. A1, P1, P2 and A2, where Ai is the support of the deck on i-abutment, while Pi is the support of the deck above i-pier. From this figure it can be extracted that the proposed interlock-ing system is able to reduce the longitudinal movements of the deck from 20% to 31%. The aforementioned reductions in the longitudinal movements of the deck also lead to reductions in the actions of the piers i.e. shear and moment actions, as the piers are monolithically connected to the deck. This improvement of the seis-mic response can lead to cost-effective bridge design, as on the one hand the piers’ reinforcement can be reduced and on the other hand the cost of the sliding bearings is reduced, as smaller bearings are adequate for the pre-assumed seismic action. Furthermore, the maintenance cost of the bridge is expected to be lower, in com-parison to the maintenance cost of the conventional bridge system, due to the elimination of the needed modular expansion joints and due to the lower cost of the smaller bearings, which would probably be replaced after their service life.


Figure 5 also shows that the proposed interlocking system is more efficient in bridges which are founded on soft ground types i.e. C instead of A, as the interlocking sys-tem lead to higher efficiency in that case - up to 10%. Specifically, the movements of the deck of the bridge, which is founded on ground type A, are up to 21% reduced, while the corresponding reduction is 31% for the case that the bridge is founded on the more flexible soil C. The increased efficiency of the proposed interlocking sys-tem in bridges, which are founded on soft ground types, can be attributed to the fact that these bridge systems respond with large seismic displacements and by extension the interlocking system is participating strongly during earthquake. Figure 6 illustrates the percentage reductions - P.R. factors - in the transverse movements of the deck above the supports. The aforementioned percentages re-sulted from the comparison of the response of the conventional and the unconven-tional bridge system, in which the proposed interlocking system was implemented. Figure 6 shows that the participation of the interlocking system leads to a 27% reduction in the transverse movement of the deck, in the unconventional bridge system in case of a relatively flexible - i.e. C - foundation soil. By contrast, if the unconventional bridge system is founded on soil types A or B then the transverse movements of its deck are increased by up to 9%. However, this consequence, which is attributed to the strong alternation of the dynamic system of the bridge when the interlocking system is activated, does not lead to an increase in the piers’ reinforcement. It is noted that in Figure 6 no P.R. factors are given above the abut-ments as the transverse displacements of the deck of the conventional and the un-conventional bridge system are restrained by stoppers over them. Figure 7 shows the variation of P.R. factor for the increased seismicity, that is ag = 0.24 g. The P.R. factor corresponds to the reduction in the longitudinal move-ments of the unconventional bridge system. From this figure it can be extracted that the longitudinal movements of the unconventional bridge deck are up to 38% re-duced. Generally, the efficiency of the proposed interlocking system is reducing while the soil type is stiffer. Specifically, the corresponding reductions in the longi-tudinal movements of the deck are up to 30% if the unconventional bridge system is founded on a stiff soil A. Similarly, the transverse movements of the deck are up to 28% reduced in the unconventional bridge system, if the bridge is founded on a soft soil (Fig. 8). The negligible increase in the transverse movements of the deck in the transverse direction of the bridge does not influence the reinforcement re-quirements of the piers. The comparison of Figures 5 and 7 leads to the conclusion that, in general, the proposed interlocking system is more efficient in bridges which are founded on areas with high seismicity, which is reflected, basically, on the design ground ac-celeration used in the analysis. The aforementioned note verifies the former re-mark, which concerns the influence of the flexibility of the soil on the efficiency of the interlocking system, which leads to the conclusion that the proposed system is more efficient in bridges which respond with large seismic displacements.


0%

10%

20%

30%

40%

50%

A1 P1 P2 A2Joint of the deck over:

P.R

. % V

aria

tion

of

long

. mov

emen

ts u

x A B CSoil:ag=0,16g

Fig. 5: The percentage reduction (P.R.) in the longitudinal movements of the deck of the

unconventional bridge system for three differ-ent soil types, (ag=0,16g).

-15%

-5%

5%

15%

25%

35%

A1 P1 P2 A2Joint of the deck over:

P.R

. % V

aria

tion

of

trans

v. m

ovem

ents

u y A B C

ag=0,16g

stoppers

Soil:

stoppers

Fig. 6: The percentage reduction (P.R.) in the transverse movements of the deck of

the unconventional bridge system for three different soil types, (ag=0,16g).

0%

10%

20%

30%

40%

50%

Α1 Μ1 Μ2 Α2Joint of the deck over:

P.R

. % V

aria

tion

of

long

. mov

emen

ts u

x A B CSoil:

ag=0,24g

Fig. 7: The percentage reduction (P.R.) in the longitudinal movements of the deck of the

unconventional bridge system for three differ-ent soil types, (ag=0,24g).

-15%

-5%

5%

15%

25%

35%

Α1 Μ1 Μ2 Α2Joint of the deck over:

P.R

. % V

aria

tion

of

trans

v. m

ovem

ents

u y A B CSoil:

stoppers stoppers

ag=0,24g

Fig. 8: The percentage reduction (P.R.) in

the transverse movements of the deck of the unconventional bridge system for three

different soil types, (ag=0,24g).

6 CONCLUSIONS

The present study proposes an interlocking system, which restrains the free seismic movement of balanced cantilever bridges lying between tunnels. This semi-connection results in the desired complex interaction of the bridge with the tunnels and by extension to the enhancement of bridge’s seismic actions. The interlocking system consists of the continuity slab, which constitutes the extension of the deck slab of the bridge towards the tunnels, and the linking-key, which offers the semi-connection of the continuity slab with the appropriately reformed foundation of the


tunnel. A parametric study performed in order to determine the efficiency of the interlocking system, in terms of reductions in the seismic displacements of the deck, came up to the following conclusions¨ 1) The proposed interlocking system can be implemented in all reinforced concrete bridge structures. Its construction is considered to be relatively easy and its cost is low, compared with the cost of the total bridge structure. The resulting bridge deck is continuous and by extension the maintenance cost of the bridge is expected to be reduced. 2) The proposed inter-locking system is possible to accommodate the serviceability requirements of the bridge deck, through the in-service allowable cracking of the connecting slab and the provision of expansion joints at the linking-key. It is noted that the minimization of the widths of the joints leads to an increased seismic par-ticipation of the interlocking system and by extension to the efficient enhancement of the seismic response of the bridge. 3) The proposed interlocking system is possible to reduce the longitudinal seismic movements of the deck from 21% to 38%. The aforementioned reductions lead to equal reductions in the seismic actions of the piers and their foundations. It is worth noting that also the dimensions of the required bearings are reduced, as the selection of these elements is based on the displacements of the deck. 4) The proposed interlocking of the bridge with the tunnels is generally more effi-cient in bridge structures which respond with large displacements, due to the in-crease in the seismic participation of the proposed system. By extension, bridges which are founded on flexible soils and bridges built on areas with high seismicity can effectively develop the proposed technique. 5) In the transverse direction of the bridge, the restrain of the seismic movements of the deck by the inter-locking system, whose connecting slab is acting as a plate, lead to a reduction in the seismic displacements of the deck by up to 28%. The negligible increase in the movements of the deck in bridges founded on stiff soils does not lead to an increase in the piers’ reinforcements. 6) The study needs a complement with analysis of more cantilever bridge struc-tures, as the length of the bridge and the height of the piers were not included in the parametric study, and they are considered to be of great importance as far as con-cerns the efficiency of the proposed interlocking system.

REFERENCES

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[2] Bishara A.G., Papakonstantinou G., Analysis of cast-in-place concrete segmen-tal cantilever bridges, Journal of structural engineering, Vol. 116, No.5, 1990.

[3] Pfeil W., Twelve years monitoring of a long span prestressed concrete bridge, Concrete International, Vol. 3,Issue 08, p. 79-84, 1981.


[4] Iglesias C., Long-term behavior of precast segmental cantilever bridges, Jour-nal of bridge engineering, Vol. 11, No.3, 2006.

[5] EN 1998-2:2005, Eurocode 8: Design of structures for earthquake resistance, Part 2: Bridges, 2005.

[6] Doğangün A., Livaoğlu R., A comparative study of the design spectra defined by Eurocode 8, UBC, IBC and Turkish Earthquake Code on R/C sample build-ings, J. Seismology, 10:335–351, DOI 10.1007/s10950-006-9020-4, 2006.

[7] PCI, (2005), Precast, prestressed concrete bridges, the high performance solution, Comprehensive Bridge Design Manual, www.pci.org/publications/bridge.

[8] Arockiasamy M. and Sivakumar M., Design Implications of Creep and Shrinkage in Integral Abutment Bridges, ACI Special Publication, Vol. 227. pp.85-106, 2005.

[9] Hindley, G, Gibbons, P, Agius, M, Carr, B, Game, R and Kashani, K, Linking Past and Present, Civil Engineering, ASCE, 2004.

[10] EN 1998-1:2005 Eurocode 8: Design of structures for earthquake resistance Part 1: General rules, seismic actions and rules for buildings, 2005.

[11] Purvis P.E.R., NCHRP 319, Bridge deck joint performance, A Synthesis of Highway Practice, Project 20-5, (Topic 30-08), 1998.

[12] Gloyd S.C., Seismic Movement at Bridge Abutments, ACI International - Special Publication SP 164-15, Vol. 164, pp. 273-288, 1996.

[13] Tegos I., Sextos A., Mitoulis S., Tsitotas M., Contribution to the improvement of the seismic performance of integral bridges, 4th European Workshop on the Seismic Behaviour of Irregular and Complex Structures, Thessaloniki, Greece, 2005.

[14] Nutt R.V. and Mayes R.L., Comparison of Typical Bridge Columns Seismi-cally Designed With and Without Abutment Participation Using AASHTO Division I-A and Proposed AASHTO LRFD Provisions, Task F3-1(a), 2000.

[15] EN 1991-1-5:2003, Eurocode 1: Actions on structures - Part 1-5: General actions -Thermal actions, prEN 1991-1-5: 2003.

[16] Kappos, A. J.,. RCCOLA-90: A microcomputer program for the analysis of the inelastic response of reinforced concrete sections, Dept. of Civ. Engrg., Aristotle University of Thessaloniki, Thessaloniki, Greece, 1993.

[17] Anagnostopoulos S.A., Equivalent viscous damping for modeling inelastic impacts in earthquake pounding problems, Earthquake Engineering & Struc-tural Dynamics, Volume 33, Issue 8, Pages: 897-902, 2004.

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[19] Computers and Structures Inc. SAP 2000 Nonlinear Version 11.0.0, User’s Reference Manual, Berkeley, California, 2007.

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