EC8 Bridges

EUROPEAN STANDARD Title Page NORME EUROPÉENNE prEN 1998-2: 2003 EUROPÄISCHE NORM Draft 3 Stage 34

English version

Eurocode 8: Design of structures for earthquake resistance Part 2: Bridges

DRAFT 3 – Stage 34 January 2003

CEN

European Committee for Standardization

Comité Européen de Normalisation Europäisches Komitee für Normung

Central Secretariat: rue de Stassart 36, B1050 Brussels

CEN 2000 Copyright reserved to all CEN members

Page 2 prEN 1998-2:2003/Draft 2 April 2002

TABLE OF CONTENTS

Page

1. INTRODUCTION 14

1.1 Scope 14 1.1.1 Scope of EN1998-2 14 1.1.2 Further parts of EN1998 14

1.2 Normative References 14 1.2.1 General reference standards 14 1.2.2 Reference Codes and Standards 15 1.2.3 Additional general and other reference standards for bridges 15

1.3 Assumptions 15

1.4 Distinction between principles and application rules 15

1.5 Definitions 15 1.5.1 Terms common to all Eurocodes 15 1.5.2 Further terms used in EN1998-2 15

1.6 Symbols 16 1.6.1 General 16 1.6.2 Further symbols used in sections 2 and 3 of EN 1998-2 16 1.6.3 Further symbols used in section 4 of EN 1998-2 17 1.6.4 Further symbols used in section 5 of EN 1998-2 18 1.6.5 Further symbols used in section 6 of EN 1998-2 19 1.6.6 Further symbols used in section 7 of EN 1998-2 20

2. BASIC REQUIREMENTS AND COMPLIANCE CRITERIA 23

2.1 Design seismic event 23

2.2 Basic requirements 24 2.2.1 Non-collapse (ultimate limit state) 24 2.2.2 Minimisation of damage (serviceability limit state) 24

2.3 Compliance criteria 24 2.3.1 General 24 2.3.2 Intended seismic behaviour 24 2.3.2.1 Ductile behaviour 25 2.3.2.2 Limited ductile/essentially elastic behaviour 26 2.3.3 Resistance verifications 27 2.3.4 Capacity design 27 2.3.5 Provisions for ductility 27


2.3.5.1 General requirement 27 2.3.5.2 Structure ductility 27 2.3.5.3 Local ductility at the plastic hinges 28 2.3.5.4 Ductility verification 29 2.3.6 Connections - Control of displacements - Detailing 29 2.3.6.1 Effective stiffness - Design seismic displacement 29 2.3.6.2 Connections 31 2.3.6.3 Control of displacements - Detailing 31 2.3.7 Simplified criteria 32

2.4 Conceptual design 33

3. SEISMIC ACTION 35

3.1 Definition of the seismic action 35 3.1.1 General 35 3.1.2 Seismological aspects 35 3.1.3 Application of the components of the motion 35

3.2 Characterization of the motion at a point 35 3.2.1 General 35 3.2.2 Quantifying of the components 36 3.2.2.1 General 36 3.2.2.2 Site dependent elastic response spectrum 36 3.2.2.3 Site dependent power spectrum 37 3.2.2.4 Time-history representation 37 3.2.2.5 Site dependent design spectrum for linear analysis 38 3.2.3 Six component model 38 3.2.3.1 General 38 3.2.3.2 Separation of the components of the seismic action 38

3.3 Characterization of the spatial variability 39

4. ANALYSIS 40

4.1 Modelling 40 4.1.1 Dynamic degrees of freedom 40 4.1.2 Masses 40 4.1.3 Element stiffness 40 4.1.4 Modeling of the soil 41 4.1.5 Torsional effects 41 4.1.6 Behaviour factors for linear analysis 42 4.1.7 Vertical component of the seismic action 45 4.1.8 Regular and irregular seismic behaviour of ductile bridges 45


4.2 Methods of analysis 46 4.2.1 Linear dynamic analysis - Response spectrum method 46 4.2.1.1 Definition, field of application 46 4.2.1.2 Significant modes 46 4.2.1.3 Combination of modal responses 47 4.2.1.4 Combination of the components of seismic action 48 4.2.2 Fundamental mode method 48 4.2.2.1 Definition 48 4.2.2.2 Field of application 48 4.2.2.3 Rigid deck model 49 4.2.2.4 Flexible deck model 50 4.2.2.5 Torsional effects in the transverse direction (rotation about the vertical axis) 50 4.2.2.6 Individual pier model 51 4.2.3 Alternative linear methods 51 4.2.3.1 Power spectrum analysis 51 4.2.3.2 Time series analysis 52 4.2.4 Non - linear dynamic time-history analysis 52 4.2.4.1 General 52 4.2.4.2 Ground motions and design combination 52 4.2.4.3 Design response effects 52 4.2.4.4 Ductile structures 53 4.2.4.5 Bridges with seismic isolation 54 4.2.5 Static non-linear analysis (pushover analysis) 54

5. STRENGTH VERIFICATION 56

5.1 Scope 56

5.2 Materials and Design strength 56 5.2.1 Materials 56 5.2.2 Design strength 56

5.3 Capacity design 56

5.4 Second order effects 58

5.5 Design seismic combination 58

5.6 Resistance verification of concrete sections 59 5.6.1 Design effects 59 5.6.2 Structures of limited ductile behaviour 59 5.6.3 Structures of ductile behaviour 59 5.6.3.1 Flexural resistance of sections of plastic hinges 59 5.6.3.2 Flexural resistance of sections outside the region of plastic hinges 60 5.6.3.3 Shear resistance of elements outside the region of plastic hinges 60 5.6.3.4 Shear resistance of plastic hinges 60 5.6.3.5 Verification of joints adjacent to plastic hinges 61


5.7 Resistance verification for steel and composite elements 66 5.7.1 Steel piers 66 5.7.1.1 General 66 5.7.1.2 Piers as moment resisting frames 66 5.7.1.3 Piers as frames with concentric bracings 66 5.7.1.4 Piers as frames with eccentric bracings 66 5.7.2 Steel or composite deck 66

5.8 Foundations 67 5.8.1 General 67 5.8.2 Design action effects 67 5.8.3 Resistance verification 67

6. DETAILING 68

6.1 General 68

6.2 Concrete piers 68 6.2.1 Confinement 68 6.2.1.1 Rectangular sections 69 6.2.1.2 Circular sections 69 6.2.1.3 Required confining reinforcement 71 6.2.1.4 Extent of confinement - Length of potential plastic hinges 71 6.2.2 Buckling of longitudinal compression reinforcement 72 6.2.3 Other rules 73 6.2.4 Hollow piers 73

6.3 Steel piers 73

6.4 Foundations 74 6.4.1 Spread foundation 74 6.4.2 Pile foundations 74

6.5 Structures of limited ductile behaviour 74 6.5.1 Verification of ductility of critical sections 74 6.5.2 Avoidance of brittle failure of specific non-ductile components 75

6.6 Bearings and seismic links 75 6.6.1 General requirements 75 6.6.2 Bearings 76 6.6.2.1 Fixed bearings 76 6.6.2.2 Moveable bearings 76 6.6.2.3 Elastomeric bearings 77 6.6.3 Seismic links, holding-down devices, shock transmission units 77 6.6.3.1 Seismic links 77 6.6.3.2 Holding-down devices 78


6.6.3.3 Shock transmission units (STU) 78 6.6.4 Minimum overlap lengths 79

6.7 Concrete abutments and retaining walls 80 6.7.1 General requirements 80 6.7.2 Abutments flexibly connected to the deck 81 6.7.3 Abutments rigidly connected to the deck 81 6.7.4 Culverts with large overburden 82 6.7.5 Retaining walls 83

7. BRIDGES WITH SEISMIC ISOLATION 84

7.1 Scope 84

7.2 Definitions and symbols 84

7.3 Basic requirements and Compliance criteria 85

7.4 Seismic action 86 7.4.1 Design spectra 86 7.4.2 Time-history representation 86

7.5 Analysis procedures and modeling 86 7.5.1 General 86 7.5.2 Design properties of the isolating system 87 7.5.2.1 Stiffness in vertical direction 87 7.5.2.2 Design properties in horizontal directions 87 7.5.2.3 Variability of properties of the isolator units 91 7.5.3 Conditions for application of analysis methods 92 7.5.4 Fundamental mode spectrum analysis 92 7.5.5 Multimode Spectrum Analysis 95 7.5.6 Time history analysis 97 7.5.7 Vertical component of seismic action 97

7.6 Verifications 97 7.6.1 Design seismic combination 97 7.6.2 Isolating system 97 7.6.3 Substructures and superstructure 98

7.7 Special requirements for the isolating system 99 7.7.1 Lateral restoring force 99 7.7.2 Lateral restraint at the isolation interface 99 7.7.3 Inspection and Maintenance 99


Page

ANNEXES ANNEX A CHARACTERISTIC SEISMIC EVENT FOR BRIDGES AND RECOMMENDATIONS FOR THE SELECTION OF DESIGN

SEISMIC ACTION DURING THE CONSTRUCTION PHASE (Informative) 99

A.1 Characteristic seismic event 99 A.2 Design seismic action for the construction phase 99 ANNEX B RELATIONSHIP BETWEEN DISPLACEMENT DUCTILITY AND CURVATURE DUCTILITY OF PLASTIC HINGES IN CONCRETE PIERS (Informative) 100 ANNEX C ESTIMATION OF THE EFFECTIVE STIFFNESS OF R. CONCRETE DUCTILE MEMBERS (Informative) 101 C.1 General 101 C.2 Method 1 101 C.3 Method 2 101 ANNEX D SPATIAL VARIABILITY AND ROTATIONAL COMPONENTS OF EARTHQUAKE MOTION (Informative) 103 D.1 General 103 D.2 Variability of earthquake motion 103 D.2.1 Introduction 103 D.2.2 Wave propagation 103 D.2.3 Simplified model 105 D.2.3.1 Fundamentals 105 D.2.3.2 Response spectra 105 D.2.3.3 Power spectra 107 D.2.3.4 Time history representation 107 D.2.4 Relative static displacements model 107 D.2.4.1 General 107 D.2.4.2 Relative static displacements 107 D.2.4.3 Design action effects 108 D.3 Rotational components 108 D.3.1 Introduction 108 D.3.2 Wave propagation 109 D.3.3 Response spectra 110


Page

D.3.4 Power spectra 111 D.3.5 Time history representation 111 ANNEX E PROBABLE MATERIAL PROPERTIES AND PLASTIC

HINGE DEFORMATION CAPACITIES FOR NON-LINEAR ANALYSES (Normative) 112

E.1 Scope 112 E.2 Probable material properties 112 E.2.1 Concrete 112 E.2.2 Reinforcement steel 114 E.2.3 Structural steel 114 E.3 Rotation capacity of plastic hinges 115 E.3.1 General 115 E.3.2 Reinforced concrete 115 ANNEX F ADDED MASS OF ENTRAINED WATER FOR IMMERSED PIERS (Normative) 118 ANNEX G CALCULATION OF CAPACITY DESIGN EFFECTS (Normative) 119 G.1 General procedure 119 G.2 Simplifications 119 ANNEX H ANALYSIS OF IRREGULAR BRIDGES (Normative) 121 H.1 General 121 H.2 Application of static non-linear analysis (pushover) 121 H.2.1 Analysis directions, reference point and target displacements 121 H.2.2 Load distribution 122 H.2.3 Deformation demands 123 H.2.4 Deck verification 123 H.2.5 Verification against non-ductile failure modes and soil failure 123 ANNEX J TESTS AND VARIATION OF DESIGN PROPERTIES OF

SEISMIC ISOLATOR UNITS (Normative) 124 J.1 Scope 124 J.2 Prototype tests 124 J.2.1 General 124


Page J.2.2 Sequence of tests 124 J.2.3 Determination of isolators characteristics 125 J.2.3.1 Force-displacement characteristics 125 J.2.3.2 Damping characteristics 126 J.2.3.3 System adequacy 126 J.3 Other tests 127 J.3.1 Wear and fatigue tests 127 J.3.2 Low temperature test 127 J.4 Verification of properties of isolating units 127 J.4.1 General 127 J.4.2 λmax-values for elastomeric bearings 129 J.4.3 λmax-values for sliding isolator units 130 ANNEX K VERIFICATION OF ELASTOMERIC BEARINGS UNDER

SEISMIC DESIGN SITUATIONS (Normative) 132 K.1 General 132 K.2 Components of shear strain 132 K.2.1 Total design shear strain 132 K.2.2 Shear strain due to compression 132 K.2.3 Shear strain due to the total seismic design displacement 133 K.2.4 Shear strain due to angular rotations 134 K.3 Design criteria for elastomeric bearings 134 K.3.1 General 134 K.3.2 Maximum values of shear strains 134 K.3.3 Stability 135 K.3.4 Fixing of bearings 135


Foreword This European Standard EN 1998-2, Eurocode 8: Design of structures for earthquake resistance. Part 2 – Bridges, has been prepared on behalf of Technical Committee CEN/TC250 «Structural Eurocodes», the Secretariat of which is held by BSI. CEN/TC250 is responsible for all Structural Eurocodes. The text of the draft standard was submitted to the formal vote and was approved by CEN as EN 1998-2 on YYYY-MM-DD. No existing European Standard is superseded. Background of the Eurocode programme In 1975, the Commission of the European Community decided on an action programme in the field of construction, based on article 95 of the Treaty. The objective of the programme was the elimination of technical obstacles to trade and the harmonisation of technical specifications. Within this action programme, the Commission took the initiative to establish a set of harmonised technical rules for the design of construction works which, in a first stage, would serve as an alternative to the national rules in force in the Member States and, ultimately, would replace them. For fifteen years, the Commission, with the help of a Steering Committee with Representatives of Member States, conducted the development of the Eurocodes programme, which led to the first generation of European codes in the 1980s. In 1989, the Commission and the Member States of the EU and EFTA decided, on the basis of an agreement1 between the Commission and CEN, to transfer the preparation and the publication of the Eurocodes to CEN through a series of Mandates, in order to provide them with a future status of European Standard (EN). This links de facto the Eurocodes with the provisions of all the Council’s Directives and/or Commission’s Decisions dealing with European standards (e.g. the Council Directive 89/106/EEC on construction products - CPD - and Council Directives 93/37/EEC, 92/50/EEC and 89/440/EEC on public works and services and equivalent EFTA Directives initiated in pursuit of setting up the internal market). The Structural Eurocode programme comprises the following standards generally consisting of a number of Parts: 1990 Eurocode : Basis of Structural Design 1991 Eurocode 1: Actions on structures 1992 Eurocode 2: Design of concrete structures 1993 Eurocode 3: Design of steel structures 1994 Eurocode 4: Design of composite steel and concrete structures

1 Agreement between the Commission of the European Communities and the European Committee for Standardisation

(CEN) concerning the work on EUROCODES for the design of building and civil engineering works (BC/CEN/03/89).


1995 Eurocode 5: Design of timber structures 1996 Eurocode 6: Design of masonry structures 1997 Eurocode 7: Geotechnical design 1998 Eurocode 8: Design of structures for earthquake resistance 1999 Eurocode 9: Design of aluminium structures Eurocode standards recognise the responsibility of regulatory authorities in each Member State and have safeguarded their right to determine values related to regulatory safety matters at national level where these continue to vary from State to State. Status and field of application of Eurocodes The Member States of the EU and EFTA recognise that Eurocodes serve as reference documents for the following purposes: - as a means to prove compliance of building and civil engineering works with the essential

requirements of Council Directive 89/106/EEC, particularly Essential Requirement N°1 – Mechanical resistance and stability – and Essential Requirement N°2 – Safety in case of fire;

- as a basis for specifying contracts for construction works and related engineering services ; - as a framework for drawing up harmonised technical specifications for construction products

(ENs and ETAs) The Eurocodes, as far as they concern the construction works themselves, have a direct relationship with the Interpretative Documents2 referred to in Article 12 of the CPD, although they are of a different nature from harmonised product standards3. Therefore, technical aspects arising from the Eurocodes work need to be adequately considered by CEN Technical Committees and/or EOTA Working Groups working on product standards with a view to achieving full compatibility of these technical specifications with the Eurocodes. The Eurocode standards provide common structural design rules for everyday use for the design of whole structures and component products of both a traditional and an innovative nature. Unusual forms of construction or design conditions are not specifically covered and additional expert consideration will be required by the designer in such cases.

2 According to Art. 3.3 of the CPD, the essential requirements (ERs) shall be given concrete form in interpretative documents for

the creation of the necessary links between the essential requirements and the mandates for harmonised ENs and ETAGs/ETAs.

3 According to Art. 12 of the CPD the interpretative documents shall :

a) give concrete form to the essential requirements by harmonising the terminology and the technical bases and indicating

classes or levels for each requirement where necessary ;

b) indicate methods of correlating these classes or levels of requirement with the technical specifications, e.g. methods of

calculation and of proof, technical rules for project design, etc.;

c) serve as a reference for the establishment of harmonised standards and guidelines for European technical approvals.


National Standards implementing Eurocodes The National Standards implementing Eurocodes will comprise the full text of the Eurocode (including any annexes), as published by CEN, which may be preceded by a National title page and National foreword, and may be followed by a National annex. The National annex may only contain information on those parameters which are left open in the Eurocode for national choice, known as Nationally Determined Parameters, to be used for the design of buildings and civil engineering works to be constructed in the country concerned, i.e.: - values and/or classes where alternatives are given in the Eurocode, - values to be used where a symbol only is given in the Eurocode, - country specific data (geographical, climatic, etc.), e.g. snow map, - the procedure to be used where alternative procedures are given in the Eurocode. It may also contain - decisions on the use of informative annexes, and - references to non-contradictory complementary information to assist the user to apply the

Eurocode. Links between Eurocodes and harmonised technical specifications (ENs and ETAs) for products There is a need for consistency between the harmonised technical specifications for construction products and the technical rules for works4. Furthermore, all the information accompanying the CE Marking of the construction products which refer to Eurocodes shall clearly mention which Nationally Determined Parameters have been taken into account. Additional information specific to EN 1998-2 The scope of this Part of EN 1998 is defined in 1.1. Except if otherwise mentioned in this Part the seismic actions are as defined in Section 3 of EN 1998-1. Due to the peculiarities of the bridge seismic resisting systems, in comparison to those of buildings and other structures, all other sections of this Part are in general not directly related to those of EN 1998-1. However several provisions of EN 1998-1 as used by direct reference.

The Eurocodes, de facto, play a similar role in the field of the ER 1 and a part of ER 2.

4 see Art.3.3 and Art.12 of the CPD, as well as clauses 4.2, 4.3.1, 4.3.2 and 5.2 of ID 1.


Since the seismic action is mainly resisted by the piers and the latter are usually foreseen from reinforced concrete, a greater emphasis has been given to such piers. Bearings are in many cases important parts of the seismic resisting system of a bridge and are therefore correspondingly treated. The same is valid for seismic isolation devices. National annex for EN 1998-2 Reference Item National Annex or

Particular Project 2.1(3) Importance factors N.A. 2.3.6.3(5) Fractions of design displacements for non-critical structural

elements N.A.

2.3.7(1) Regions of low seismicity N.A. 2.3.7(1) Simplified criteria for bridges in regions of low seismicity N.A. 3.3(1)P Simplified models for spatial variability of seismic action N.A. or P.P. 4.1.2(4) ψ values for traffic loads to be assumed concurrent to

seismic action N.A.

5.4(1) Simplified methods for second order effects N.A. 6.5.1 (1)P Simplified verification rules for bridges in low seismicity

regions N.A.

6.6.3.2(1)P Necessity for holding-down devices N.A. J.3.2 Design low temperature for seismic situation Tmin,sd N.A.


1. INTRODUCTION

1.1 Scope 1.1.1 Scope of EN1998-2

(1) Within the framework of the scope set forth in EN 1998-1, this part of the Code contains the particular Performance Requirements, Compliance Criteria and Application Rules applicable to the earthquake resistant design of bridges. (2) This Part covers primarily the seismic design of bridges in which the horizontal seismic actions are mainly resisted through bending of the piers or at the abutments i.e. bridges composed of vertical or near vertical pier systems supporting the traffic deck superstructure. It is also applicable for the seismic design of cable stayed and arched bridges, although its provisions should not be considered as fully covering these cases. (3) Suspension bridges, timber and masonry bridges, moveable bridges and floating bridges are not included in the scope of this Part. (4) Part 2 of EN 1998 contains only those provisions that in addition to this provisions of the other relevant Eurocodes or relevant Parts of EN 1998, must be observed for the design of bridges in seismic regions. (5) The following topics are dealt with in the text and annexes of this Part. Basic requirements and Compliance Criteria, Seismic Action, Analysis, Strength Verification, Detailing. Finally, this Part also includes a special section on seismic isolation with provisions covering the application of this method of seismic protection to bridges. 1.1.2 Further parts of EN1998 See EN1998-1. 1.2 Normative References (1)P The following normative documents contain provisions, which through references in this text, constitute provisions of this European standard. For dated references, subsequent amendments to or revisions of any of these publications do not apply. However, parties to agreements based on this European standard are encouraged to investigate the possibility of applying the most recent editions of the normative documents indicated below. For undated references the latest edition of the normative document referred to applies. 1.2.1 General reference standards See EN1998-1.


1.2.2 Reference Codes and Standards See EN 1998-1. 1.2.3 Additional general and other reference standards for bridges

EN 1990: Annex 2 Basis of structural design: Application for bridges EN 1991-2:200x Actions on structures: Traffic loads on bridges EN 1992-2:200x Design of concrete structures. Part 2 – Bridges EN 1993-2:200x Design of steel structures. Part 2 – Bridges EN 1998-1:200x Design of structures for earthquake resistance. General rules, seismic

actions and rules for buildings EN 1998-5:200x Design of structures for earthquake resistance. Foundations, retaining

structures and geotechnical aspects.

1.3 Assumptions (1) In addition to the general assumptions of clause 1.3 of EN 1990:2002, the following assumption applies. (2)P It is assumed that no change of the structure will take place during the construction phase or during the subsequent life of the structure, unless proper justification and verification is provided. Due to the specific nature of the seismic response this applies even in the case of changes that lead to an increase of the structural resistance. 1.4 Distinction between principles and application rules (1) The rules of clause 1.4 of EN 1990:2002 apply. 1.5 Definitions

1.5.1 Terms common to all Eurocodes (1) The terms and definitions of clause 1.5 of EN 1990:2002 apply. 1.5.2 Further terms used in EN1998-2 Capacity design: The design procedure used in structures of ductile behaviour to secure the hierarchy of strengths of the various structural components necessary for leading to the intended configuration of plastic hinges and for avoiding brittle failure modes. Ductile elements: Elements able to dissipate energy through the formation of plastic hinges.


Ductile structure behaviour: Structures that under strong seismic motions can dissipate significant amount of input energy through the formation of an intended configuration of plastic hinges or by other mechanisms. Seismic Isolation: The provision of bridge structures with special isolating devices for the purpose of reducing the seismic response (forces and/or displacements). Spatial variability: Spatial variability of the seismic action means that the motion at different supports of the bridge is assumed to be different and, as a result, the definition of the seismic action cannot be based on the characterization of the motion at a single point, as is usually the case. Seismic behaviour: The behaviour of the bridge under the design seismic event which, depending on the characteristics of the global force-displacement relationship of the structure, can be ductile or limited ductile/essentially elastic. Seismic links: Restrainers through which part or all of the seismic action may be transmitted. Used in combination with bearings, they may be provided with appropriate slack so as to be activated only in the case when the design seismic displacement is exceeded. Min. overlap length: A safety measure in the form of a minimum distance between the inner edge of the supported and the outer edge of the supporting element. The minimum overlap is intended to ensure that the function of the support is maintained under extreme seismic displacements. Design seismic displacement: The displacement induced by the design seismic actions. Total design displacement under seismic conditions: The displacement used to determine adequate clearances for the protection of critical or major structural elements. It includes the design seismic displacement, the displacement due to the long term effect of the permanent and quasi-permanent actions and an appropriate fraction of the displacement due to thermal movements. 1.6 Symbols 1.6.1 General (1) The symbols indicated in clause 1.6 of EN 1990:2002 apply. For the material-dependent symbols, as well as for symbols not specifically related to earthquakes, the provisions of the relevant Eurocodes apply. (2) Further symbols, used in connection with the seismic actions, are defined in the text where they occur, for ease of use. However, in addition, the most frequently occuring symbols used in EN 1998-2 are listed and defined in the following subsections. 1.6.2 Further symbols used in sections 2 and 3 of EN 1998-2 Aed design seismic action dE design seismic displacement (due only to seismic action)


dEe seismic displacement determined from linear analysis dG displacement due to the permanent and quasi-permanent actions measured in long term dT displacement due to thermal movements du ultimate displacement dy yield displacement FRd design resisting force to the earthquake action Ri reaction force on the base of pier i Sa site averaged response spectrum Si site dependent response spectrum Teff effective period of the isolation system γI importance factor µd displacement ductility ψ2 combination factor for the quasi-permanent value of thermal action 1.6.3 Further symbols used in section 4 of EN 1998-2 AEd design seismic action AEx seismic action in direction X AEy seismic action in direction Y AEz seismic action in direction Z B width of the deck da maximum average of the displacements in the transverse direction of all pier tops under the

transverse seismic action or under the action of a transverse load of similar distibution di displacement of the i-th nodal point dm assymtotic value of the spectrum for the m-th motion for large periods, expressed in

displacements E probable maximum value of an action effect Ei modal responses e ea + ed ea accidental mass eccentricity (= 0,03L or 0,03B) ed additional eccentricity reflecting the dynamic effect of simultaneous translational and

torsional vibration (= 0,05L or 0,05B) eo theoretical eccentricity F horizontal force determined according to the fundamental mode method G total effective weight of the structure, equal to the weight of the deck plus the weight of

the upper half of the piers Gi weight concentrated at the i-th nodal point g acceleration of gravity h depth of the cross section in the direction of flexure of the plastic hinge K stiffness of the system km effect of the m-th independent motion L total length of the continuous deck L distance from the plastic hinge to the point of zero moment M total mass MEd,i maximum value of design moment under the seismic load combinations at the intended

location of plastic hinge of ductile member i MRd,i design flexural resistance of the plastic hinge section of ductile member i


Mt equivalent static moment acting about the vertical axis through the centre of mass of the

deck Qk,1 characteristic value of traffic load Rd design resistance Sd(T) spectral acceleration of the design spectrum T period of the fundamental mode of vibration for the direction under consideration X horizontal longitudinal axis Y horizontal transverse axis Z vertical axis ri required local force reduction factor at ductile member i rmin minimum (ri) rmax maximum (ri) αs shear ratio of the pier ∆d maximum difference of the displacements in the transverse direction of all pier tops under

the transverse seismic action or under the action of a transverse load of similar distibution ηk normalized axial force (= NEd/(Acfck)) θp,d design plastic rotation capacity θp,E plastic hinge rotation demand ξ the viscous damping ratio ψ2,1 quasi-permanent values of combination factor of variable action 1.6.4 Further symbols used in section 5 of EN 1998-2 Ac area of the concrete section AEd seismic design action effect (seismic action alone) ASd design seismic combination Asx area of horizontal joint reinforcement dEd relative transverse displacement of the ends of the ductile member under consideration Ed design value of action effect of seismic combination FC capacity design effect fck characteristic concrete strength fctd design tensile strength of concrete fsd reduced stress of reinforcement, for reasons of limitation of cracking fsy design yield strength of the joint reinforcement Gk permanent load with their characteristic values Mo overstrength moment MEd design moment under the seismic load combination MRd design flexural strength of the section NEd axial force corresponding to the design seismic combination NcG axial force of column under permanent actions Njz vertical axial joint force Q1k characteristic value of the traffic load Q2 quasi permanent value of actions of long duration Pk characteristic value of prestressing after all losses Rd design resistance of the section Rdf maximum design friction force of sliding bearing TRc resultant force of the tensile reinforcement of the column VE,d design shear force Vjx design horizontal shear of the joint


Vjz design vertical shear of the joint V1bC shear force of the beam adjacent to the tensile face of the column zb internal force lever arm of the beam end sections zc internal force lever arm of the plastic hinge section of column γM material safety factor γo overstrength factor γof magnification factor for friction due to ageing effects ρx horizontal reinforcement ratio ρy reinforcement ratio of closed stirrups in the transverse direction of the joint panel

(orthogonal to the plane of action) ρz vertical reinforcement ratio ψ21 combination factor 1.6.5 Further symbols used in section 6 of EN 1998-2 Ac gross concrete area of the section Acc confined (core) concrete area of the section ag ground acceleration as defined in 3.2.2.2 of EN 1998-1 Asp area of the spiral or hoop bar Asw total area of hoops or ties in the one direction of confinement At area of one tie leg b dimension of the concrete core perpendicular to the direction of the confinement

under consideration, measured to the outside of the perimeter hoop bmin smallest dimension of the concrete core ca apparent phase velocity of the seismic waves in the soil deg effective displacement of the two parts due to differential seismic ground

displacement des effective seismic displacement of the support due to the deformation of the structure dg design value of the peak ground displacement as defined by 3.2.2.4 of EN 1998-1 Di inside diameter Dsp diameter of the spiral or hoop bar Ed total earth pressure acting on the abutment under seismic conditions as per EN1998-5 FRd design resistance ft tensile strength fy yield strength fys yield strength of the longitudinal reinforcement fyt yield strength of the tie Lh design length of plastic hinges Leff effective length of deck lm minimum support length securing the safe transmission of the vertical reaction loν minimum overlap length Qd the weight of the section of the deck linked to a pier or abutment, or the least of the

weights of the two deck sections on either side of the intermediate separation joint S soil parameter as defined in 3.2.2.2 of EN 1998-1 s distance between tie legs sL maximum (longitudinal) spacing sT distance between hoop legs or supplementary cross ties st transverse (longitudinal) spacing


TC elastic spectrum parameter as defined in 3.2.2.2 of EN 1998-1 vg peak ground velocity vs shear wave velocity in the soil under the shear strain αg design ground acceleration γI importance factor γs free-field seismic shear deformation of the soil δ parameter depending on the ratio ft/fy µΦ required curvature ductility ΣAs sum of the areas of the longitudinal bars restrained by the tie ρL reinforcement ratio of the longitudinal reinforcement ρw transverse reinforcement ratio φL diameter of the longitudinal bar 1.6.6 Further symbols used in section 7 of EN 1998-2 ag design ground acceleration on rock corresponding to the importance category of the

bridge ag,475 design ground acceleration corresponding to a design seismic event with reference

return period 475 years ddc design displacement of the effective stiffness center in the direction considered ddi total design displacement of isolator unit i in the direction considered dd,f design displacement of the stiffness center derived by the fundamental mode method dd,m resulting displacement of the stiffness center derived from the analysis dmax maximum displacement dy yield displacement ED dissipated energy per cycle at the design displacement ddc EDi dissipated energy per cycle of isolator unit i, at the design displacement ddc EE design seismic forces EEA seismic internal forces derived from the analysis etot,x = eacc+ ex total eccentricity in the longitudinal direction ex eccentricity between stiffness center and center of mass in the longitudinal direction Fy yield force at monotonic loading F0 force at zero displacement at cyclic loading G shear modulus i isolator unit Ke elastic stiffness at monotonic loading Kp post elastic (tangent) stiffness Keff effective stiffness of the isolation system in the principal horizontal direction under

consideration, at a displacement equal to the design displacement ddc Keff,i composite stiffness of isolator units and the corresponding pier i Kf,i rotation stiffness of foundation of pier i Ki,i effective stiffness Ks,i displacement stiffness Kyi effective composite stiffnesses of isolator unit and pier i, in the y direction


L total length of the deck between joints NSd normal force through the device r radius of gyration of the deck mass about vertical axis through its center of mass S shape factor

sign(•

d ) sign of the velocity vector •

d S, TC, TD parameters of the design spectrum depending on the soil, according to 3.2.2.2 of

EN1998-1 Teff effective period v velocity of relative motion at the interface of a viscous isolator Vd maximum shear force Vd,f shear force transferred through the isolation interface derived by the fundamental

mode method Vd,m total shear force transferred through the isolation interface derived from the analysis Wd weight of the superstructure mass xi coordinate of isolator unit i relative to the effective stiffness center UBDP Upper bound design properties of isolators LBDP Lower bound design properties of isolators α exponent of velocity in viscous dampers γI importance factor of the bridge µd dynamic friction coefficient ξ equivalent viscous damping ratio ξeff effective damping of the isolation system



2. BASIC REQUIREMENTS AND COMPLIANCE CRITERIA

2.1 Design seismic event (1)P The design philosophy of this Code, regarding the seismic resistance of bridges, is based on the general requirement that emergency communications shall be maintained, with appropriate reliability, after the design seismic event (AEd). (2)P Unless otherwise specified in this part, the elastic spectrum of the characteristic seismic action shall be taken from 3.2.2.2, 3.2.2.3 and 3.2.2.4 of EN 1998-1. For application of the equivalent linear method of 4.1.6 (using behaviour factor q) the spectrum shall be taken from 3.2.2.5 of EN 1998-1.

Note: Annex A gives information on the characteristic seismic action and on the selection of the design seismic action during the construction phase.

(3) Differentiation of target reliability may, in the absence of reliable statistical evaluation of seismological data, be implemented by defining the design seismic action (AEd) as : AEd = γIAEk (2.1) where γI is the importance factor. Note: The values of importance factors may be set by the National Annex. Recommended values are as

follows: Bridge Importance Category Importance Factor γI Greater than average 1,30 Average 1,00 Less than average 0,70 (4) In general bridges on motorways and national roads as well as railroad bridges are considered to belong to the category of “average” importance.

(5) To the category of "greater than average" importance belong bridges of critical importance for maintaining communications, especially after a disaster, bridges whose failure is associated with a large number of probable fatalities, and major bridges for which a design life greater than normal is required. (6) To the category of "less than average" importance belong bridges which are not critical for communications and for which the adoption of either the standard probability of exceedence of the design seismic event or the normal bridge design life, is not economically justifiable. (7) Recommendations for the selection of the design seismic event, appropriate for use during the construction period of bridges are given in Annex A.


2.2 Basic requirements (1)P The design shall aim to fulfil the following two basic requirements. 2.2.1 Non-collapse (ultimate limit state) (1)P After the occurrence of the design seismic event, the bridge shall retain its structural integrity and adequate residual resistance, although at some parts of the bridge considerable damage may occur. (2)P The bridge shall be damage-tolerant, i.e. those parts of the bridge susceptible to damage by their contribution to energy dissipation during the design seismic event shall be designed in such a manner as to ensure that, following the seismic event, the structure can sustain the actions from emergency traffic, and inspections and repair can be performed easily. (3)P To this end, flexural yielding of specific sections (i.e. the formation of plastic hinges) is allowed in the piers, and is in general necessary, in regions of high seismicity, in order to reduce the design seismic action to a level requiring reasonable additional construction costs. (4)P The bridge deck shall in general be designed to avoid damage, other than local to secondary elements such as expansion joints, continuity slabs (see 2.3.2.1(4)) or parapets. It must therefore, be able to sustain the loads from piers undergoing plastic hingeing and must not become unseated under extreme seismic displacements. 2.2.2 Minimisation of damage (serviceability limit state) (1)P Earthquakes with high probability of occurrence during the design life of the bridge, should cause only minor damage to secondary elements and those parts of the bridge intended to contribute to energy dissipation. There should be no need to reduce traffic over the bridge nor to carry out immediate repairs. 2.3 Compliance criteria 2.3.1 General (1)P In order to satisfy the basic requirements set forth in 2.2, the design must comply with the criteria outlined in the following Clauses. In general the criteria while aiming explicitly at satisfying the non-collapse requirement (2.2.1), implicitly cover the damage minimization requirement (2.2.2) as well. (2)P The compliance criteria depend on the behaviour which is intended for the bridge under the design seismic action. This behaviour may be selected according to 2.3.2. 2.3.2 Intended seismic behaviour (1)P The bridge shall be designed so that its behaviour under the design seismic action is either ductile, or limited ductile/essentially elastic, depending on the seismicity of the site, whether isolation technology is being adopted, or any other constraints which may prevail. This behaviour


(ductile or limited ductile) is characterized by the global force-displacement relationship of the structure (see Fig. 2.1)

1.0

1.5

3.0

BEHAVIOUR

Ideal elastic Essentially elastic

Limited ductile

Ductile

DISPLACEMENT

FORCE q

Figure 2.1 Seismic behaviour

2.3.2.1 Ductile behaviour

(1) In regions of moderate to high seismicity it is usually preferable, both for economic and safety reasons, to design a bridge for ductile behaviour i.e. to provide it with reliable means to dissipate a significant amount of the input energy under severe earthquakes. This is accomplished by providing for the formation of an intended configuration of flexural plastic hinges or by using isolating devices according to Section 7. The part of this subclause that follows refers to ductile behaviour achieved by flexural plastic hinges. (2)P Bridges of ductile behaviours shall be designed so that a dependably stable plastic mechanism can form in the structure through the formation of flexural plastic hinges, normally in the piers, which act as the primary energy dissipating components. (3) As far as possible the location of plastic hinges should be selected at points accessible for inspection and repair.


(4) The bridge deck shall remain within the elastic range. However formation of plastic hinges is allowed in flexible ductile slabs, providing top slab continuity between adjacent simply supported precast concrete girder spans. (5)P The formation of plastic hinges is not allowed in reinforced concrete sections where the normalized axial force ηk defined in 5.3.(4)P exceeds 0,6. (6) The global force-displacement relationship should show a significant force plateau at yield and should be reversible in order to ensure hysteretic energy dissipation at least over 5 deformation cycles (see Figures 2.1, 2.2 and 2.3).

Note: Elastomeric bearings used over some supports may cause some increase of the resisting force, with increasing displacements, after plastic hinges have formed in the other supporting elements. However the rate of increase of the resisting force should be appreciably reduced after the formation of plastic hinges.

(7) Flexural hinges need not necessarily form in all piers. However the optimum post-elastic seismic behaviour of a bridge is achieved if plastic hinges develop approximately simultaneously in as many piers as possible. (8) Supporting elements (piers or abutments) connected to the deck through sliding or flexible mountings (sliding bearings or flexible elastomeric bearings) should, in general, remain within the elastic range. (9) The capability of the structure to form flexural hinges is necessary in order to ensure energy dissipation and consequently ductile behaviour (see 4.1.6 (2)).

Note: The deformation of bridges supported by normal elastomeric bearings is predominantly elastic and does not lead to ductile behaviour. (see 4.1.6 (11)).

2.3.2.2 Limited ductile/essentially elastic behaviour

(1) No significant yield appears under the design earthquake. In terms of force-displacement characteristics, the formation of a force plateau is not required, while deviation from the ideal elastic behaviour provides some hysteretic energy dissipation. Such a behaviour corresponds to a behaviour factor q ≤ 1,5 and shall be referred to, in this Code, as "limited ductile".

Note: Values of q in the range 1 ≤ q ≤ 1,5 are mainly attributed to inherent margin between design and actual strength under the design seismic situations.

(2) For bridges where the seismic response may be dominated by higher mode effects (e.g cable-stayed bridges) or where the detailing for ductility of plastic hinges may not be reliable (e.g. due to the presence of high axial force or of a low shear ratio), it is preferable to select an elastic behaviour (q = 1).


2.3.3 Resistance verifications (1)P In bridges of ductile behaviour the regions of plastic hinges shall be verified to have adequate flexural strength to resist the design seismic effects as defined in 5.5. The shear resistance of the plastic hinges as well as both the shear and flexural resistances of all other regions shall be designed to resist the "capacity design effects" defined in 2.3.4 below (see also 5.3). (2) In bridges of limited ductile behaviour all sections should be verified to have adequate strength to resist the design seismic effects of 5.5 (see also 5.6.2). 2.3.4 Capacity design (1)P For bridges of ductile behaviour, capacity design shall be used to secure the hierarchy of resistances of the various structural components necessary for leading to the intended configuration of plastic hinges and for avoiding brittle failure modes. (2) This shall be achieved by designing all elements intended to remain elastic or against all brittle modes of failure, using "capacity design effects". Such effects result from equilibrium conditions at the intended plastic mechanism, when all flexural hinges have developed an upper fractile of their flexural resistance (overstrength), as defined in 5.3. (3) For bridges of limited ductile behaviour the application of the capacity design procedure is not compulsory. 2.3.5 Provisions for ductility 2.3.5.1 General requirement (1)P The intended plastic hinges shall be provided with adequate ductility to ensure the required overall structure ductility.

Note: The definitions of structure and local ductilities, given in 2.3.5.2 and 2.3.5.3 that follow, are intended to provide the theoretical basis of ductile behaviour. In general they are not required for practical verification of ductility, which is effected according to 2.3.5.4.

2.3.5.2 Structure ductility (1) Referring to an equivalent one degree of freedom system, having an idealized bilinear force-displacement relationship, as shown in Figure 2.2, the design value of the ductility of the structure (available displacement ductility) is defined as the ratio of the ultimate limit state displacement (du) to the yield displacement (dy), both measured at the centre of mass, i.e. yud /ddµ = . (2) When an equivalent linear analysis is performed, the yield force of the global bilinear diagram is assumed equal to the design resisting force FRd. The yield displacement defining the elastic branch is selected so as to best approximate the design curve (for monotonic loading).


(3) The ultimate displacement du is defined as the maximum displacement satisfying the following condition. The structure shall be capable of sustaining at least 5 full cycles of deformation to the ultimate displacement, • without initiation of failure of the confining reinforcement for reinforced concrete sections, or

local buckling effects for steel sections, and • without drop of the resisting force for steel ductile elements or without a drop exceeding 20%

of the maximum for reinforced concrete ductile elements (see Figure 2.3).

F Rd

d y d u

Elasto-plastic

Design

F Rd

d u d y

Monotonic Loading

5th cycle

Rd 0.2F < -

Figure 2.2 Figure 2.3 Global force-displacement diagram Force-displacement cycles (Monotonic loading) (Reinforced concrete)

2.3.5.3 Local ductility at the plastic hinges

(1) The structure ductility depends on the available local ductility at the plastic hinges (See Figure 2.4). This can be expressed as curvature ductility of the cross-section:

yuΦ/ΦΦµ = (2.2)

or as chord rotation ductility of the hinge, that depends on the plastic rotation capacity θp,u=θu-θy of the hinge section: yup,yyuyuθ /( θθ)/θθθ/θθµ +=−+== 11 (2.3) The chord rotation is measured over the length L, between the plastic hinge and the section of zero moment, as shown in Figure 2.4. The relation between θp, Φu, Φy, L and the length of plastic hinge Lp, is given in Annex E.

L

Lp

θ

Figure 2.4 : Chord rotation θ

M Plastic hinge

≤ 0.2Fmax


(2) In the above expressions the ultimate deformations should comply to the conditions of 2.3.5.2 (3).

Note: The relation between curvature ductility of a plastic hinge and the structure displacement ductility is given in Annex B, for a simple case. This relation is not intended for ductility verification.

2.3.5.4 Ductility verification

(1)P Compliance with the Specific Rules given in Section 6 shall be deemed, in general, to ensure the availability of adequate local and overall structure ductility. (2)P When non-linear dynamic analysis is performed, rotation demands shall be checked against available rotation capacities of the plastic hinges (see 4.2.4.4). (3) For bridges of limited ductile behaviour the provisions of 6.5 should be applied. 2.3.6 Connections - Control of displacements - Detailing

2.3.6.1 Effective stiffness - Design seismic displacement

(1)P Within the framework of equivalent linear analysis methods allowed by the present code, the stiffness of each element shall adequately approximate its deformation under the maximum stresses induced by the design seismic action. For elements containing plastic hinges this corresponds to the secant stiffness at the theoretical yield point (See Figure 2.5).

Figure 2.5

Moment - deformation diagrams at plastic hinges

Moment-Rotation of plastic hinge Moment-Curvature of cross-section Structural Steel Reinforced Concrete

Secant Stiffness

C

Secant Stiffness

C C y u

M Rd M Rd

φ

φy φy

Φy Φu Φ

MRd

θy θu θ

MRd


(2) For reinforced concrete elements in bridges of ductile behaviour, in the absence of a more rigorous method, the effective flexural stiffness to be used in linear seismic analysis (static or dynamic) may be estimated as follows: • For reinforced concrete piers, a value calculated on the basis of the secant stiffness at the

theoretical yield point. • For prestressed or reinforced concrete decks the stiffness of the uncracked sections.

Note: Annex C gives guidance for the estimation of the effective stiffness of reinforced concrete members.

(3) In limited ductile bridges the flexural stiffness of the uncracked sections may be used globally. (4) For both ductile and limited ductile bridges, the significant reduction of the torsional stiffness of concrete decks, in relation to the uncracked stiffness, should be accounted for. In the absence of a more accurate assessment the following fractions of the uncracked stiffness may be used: • for open sections or slabs, the torsional stiffness may be ignored • for prestressed box sections 50% of the uncracked stiffness • for reinforced concrete box sections 30% of the uncracked stiffness (5) Displacements resulting from an analysis according to the rules given in paragraphs (2) and (3) above, should be corrected on the basis of the flexural stiffness of the elements, corresponding to the resulting level of the actual stresses, even for limited ductile structures.

Note: It is noted that in the case of eqivalent linear analysis (see 4.3.6 (1)P) an overestimation of the effective stiffness leads to results which are on the safe side regarding the seismic actions. In such a case, only the displacements need be corrected, after the analysis, on the basis of the resulting level of moments. On the other hand, if the initial assumption of effective stiffness is significantly lower than that corresponding to the actual stresses, the analysis must be repeated using a better approximation of the effective stiffness.

(6)P The displacements dEe determined from linear seismic analysis, static or dynamic, shall be multiplied by the displacement ductility µd, to obtain the design seismic displacements dE. dE = ± µddEe (2.4)

(7)P The displacement ductility shall be assumed as follows: • when the fundamental period T in the direction under examination is T ≥ To = 1,25Tc, where

Tc has the values defined by 3.2.2.2 of EN1998-1, then

µd = q (2.5) • when T < To, then


( ) 45q1 od −≤+−= 1

TT

qµ (2.6)

where q is the value of the behaviour factor assumed in the analysis for dEe.

Note: Relation (2.6) provides a smooth transition between the “equal displacement” rule that is applicable for T ≥ To, and the short period range (not typical to bridges) where the assumption of a low q-value is expedient. For very small periods (T < 0,033s), q = 1 should be assumed (see also 4.1.6 (9)) resulting to µd = 1.

(8)P When non-linear time domain analysis is used, the deformation characteristics of the yielding elements shall adequately approximate their actual post-elastic behaviour, both in the loading and unloading branches of the hysteresis loops, as well as potential degradation effects (see 4.2.4.4).

2.3.6.2 Connections

(1)P Connections between supporting and supported elements shall be adequately designed in order to secure structural integrity and avoid unseating under extreme seismic displacements. (2)P Bearings, links and holding-down devices used for securing structural integrity, shall be designed using capacity design effects (see 5.3, 6.6.2.1, 6.6.3.1 and 6.6.3.2). (3) Appropriate overlap lengths shall be provided between supporting and supported elements at moveable connections, in order to avoid unseating (see 6.6.4). (4) As an alternative to the provision of overlap length, positive linkage between supporting and supported elements may be used (see 6.6.3).

2.3.6.3 Control of displacements - Detailing

(1)P In addition to ensuring a satisfactory overall ductility, structural and non-structural detailing must ensure adequate behaviour of the bridge and its components under the design seismic displacements. (2)P Adequate clearances shall be provided for protection of critical or major structural elements. Such clearances shall accommodate the total design value of the displacement under seismic conditions dEd determined as follows: dEd = dE + dG + ψ2dT (2.7) where, following displacements shall be combined with the most onerous sign: dE is the design seismic displacement according to equation (2.4) dG is the displacement due to the permanent and quasi-permanent actions measured in long

term (e.g. post-tensioning, shrinkage and creep for concrete decks). dT is the displacement due to thermal movements.


ψ2 is the combination factor for the quasi-permanent value of thermal action, according to Tables A2.1, A2.2 or A2.3 of Annex A2 of EN1990.

The total design seismic displacement shall be increased by the displacement due to second order effects when such effects have a significant contribution. (3) The relative seismic displacement dE between two independent sections of a bridge may be estimated as the square root of the sum of squares of the values calculated for each section. (4)P Large shock forces, caused by unpredictable impact between major structural elements, shall be prevented by means of ductile/resilient elements or special energy absorbing devices (buffers). Such elements shall have a slack at least equal to the total design displacement dEd. (5) The detailing of non-critical structural elements (e.g. deck movement joints) and abutment back-walls, expected to be damaged during the design seismic event, should cater, for a predictable mode of damage and provide for the possibility of permanent repair. Clearances should accommodate appropriate fractions of the design seismic displacement and thermal movement, after allowing for any long term creep and shrinkage effects, so that damage under frequent earthquakes can be avoided. The appropriate values of such fractions depend on techno-economical considerations.

Note 1: The National Annex may define such appropriate values to be used in the absence of an explicit optimisation. Recommended values are as follows:

40% of the design seismic displacement 50% of the thermal movement

Note 2: At joints of railway bridges, transverse differential displacement may have to be either avoided or limited to values appropriate for preventing derailments.

2.3.7 Simplified criteria (1) In regions of low seismicity, an appropriate classification of the bridges and simplified compliance criteria, pertaining to the seismic design of individual classes may be established.

Note 1: Regions of low seismicity (and by consequence those of moderate to high seismicity) may be defined by the National Annex. The criterion ag ≤ 0,10g is recommended. Note 2: Classification of bridges and simplified criteria for the seismic design pertaining to individual bridge classes in low seismicity regions may be established by the National Annex. It is recommended that these simplified criteria are based on a limited ductile/essentially elastic seismic behaviour of the bridge, for which no special ductility requirements are necessary.


2.4 Conceptual design (1) Consideration of seismic effects at the conceptual stage of the design of bridges is important even in low to moderate seismicity regions. (2) In areas of low seismicity the type of intended seismic behaviour of the bridge (see 2.3.2) must be decided. If a limited ductile (or essentially elastic) behaviour is selected, simplified criteria, in accordance with 2.3.7 may be applied. (3) In areas of moderate or high seismicity the selection of ductile behaviour is generally expedient. Its implementation, either by providing for the formation of a dependable plastic mechanism or by using base isolation and energy dissipating devices, must be decided. When a ductile behaviour is selected paragraphs (4) to (8) should be considered: (4) The number of supporting elements (piers and abutments) that will be used to resist the seismic forces in the longitudinal and the transverse direction must be decided. In general continuous structures behave better under seismic conditions than bridges having many movement joints. The optimum post-elastic seismic behaviour is achieved if plastic hinges develop approximately simultaneously in as many piers as possible. However, the number of the earthquake resisting piers may have to be reduced, by using sliding or flexible mountings between deck and piers in the longitudinal directions, to avoid high reactions due to restrained deformations. (5) A balance should be maintained between the strength and the flexibility requirements of the horizontal supports. High flexibility reduces the level of the design seismic action but increases the movement at the joints and moveable bearings and may lead to high second order effects. (6) In the case of bridges with continuous deck, in which the transverse stiffness of the abutments and of the adjacent piers is very high in relation to that of the other piers (as may occur in steep sided valleys), a very unfavourable distribution of the transverse seismic action on these elements may take place, as shown in Figure 2.6. In such cases it may be preferable to use transversally sliding or elastomeric bearings over the short piers or the abutments. (7) The location of the energy dissipating points must be chosen so as to ensure their accessibility for inspection and repair. Their locations should be made clear in the appropriate design document. (8) The location of other potential or expected damage areas under severe motions must be identified and the difficulty of repairs must be minimised. (9) In exceptionally long bridges, or in bridges crossing non-homogeneous soil formations, the number and location of intermediate movement joints must be decided. (10) In the case of bridges over potentially active tectonic faults, the probable discontinuity of the ground displacement should be estimated and accommodated either by adequate flexibility of the structure or by provision of suitable movement joints.


(11) The liquefaction potential of the foundation soil must be investigated according to the relevant provisions of EN1998-5.

Figure 2.6 Unfavourable distribution of transverse seismic action

ELEVATION

PLAN


3. SEISMIC ACTION 3.1 Definition of the seismic action 3.1.1 General (1)P The seismic action can be defined by means of different models, whose complexity shall be appropriate to the relevant earthquake motion to be described and the importance of the structure and commensurate with the sophistication of the model used for the idealization of the bridge. (2)P In this section only the shaking transmitted by the ground to the structure is considered in the quantification of the seismic action. However, earthquakes can induce permanent displacements in soils (ruptures, liquefaction of sandy layers and ground offset due to active faulting) that may result in imposed deformations with severe consequences to bridges. This type of hazard shall be evaluated through specific studies and its consequences shall be minimised by an appropriate selection of the structural foundation system and possibly ground improvement. Tsunami effects are not treated in this Code. 3.1.2 Seismological aspects (1)P In the definition of the seismic action the following aspects shall be considered: • the characterization of the motion at a point; • the characterization of the spatial variability of the motion. 3.1.3 Application of the components of the motion 1(P) In general only the three translational components of the seismic action are taken into account. When the Response Spectrum method is applied the bridge may be analyzed separately for shaking in the longitudinal, transverse and vertical directions. In this case the seismic action is represented by three one-component actions, one for each direction, quantified according to 3.2.2 and 3.2.3.2. The action effects shall be combined according to 4.2.1.4; 2(P) When non-linear time-history analysis is performed or when the six component model or the spatial variability of the seismic motion is taken into account, the bridge shall be analyzed under the simultaneous action of the different components.

3.2 Characterization of the motion at a point 3.2.1 General (1)P The characterization of the motion at a point shall be carried out in two phases: • Quantification of each component of the motion; • Construction of a three component model of the motion with three translational components,

or of a six component model of the motion, with three translational components and three rotational components.


(2) The seismic action is applied at the interface between the structure and the ground. If springs are used to represent the soil stiffness either in connection with spread footings or with deep foundations (piles, shafts etc., see Part 5) the motion is applied at the soil end of the springs. In general it is not necessary to use the three rotational components of the ground motion. If their inclusion is considered necessary, then cl. 3.2.3 is applicable. 3.2.2 Quantifying of the components

3.2.2.1 General

(1)P Each component of the earthquake motion shall be quantified in terms of a response spectrum, or a power spectrum, or a time history representation (mutually consistent) as set out in Section 3 of Part 1, which also provides the basic definitions.

3.2.2.2 Site dependent elastic response spectrum

3.2.2.2.1 Horizontal component (1)P The horizontal component shall be in accordance with 3.2.2.2 of Part 1. 3.2.2.2.2 Vertical component (1)P When needed (see 4.1.7), the site dependent response spectrum for the vertical component of the earthquake motion shall be taken in accordance with 3.2.2.3 of EN1998-1. 3.2.2.2.3 Site averaged response spectrum (1)P In the case of bridges whose abutments and piers are supported on soils having significantly different soil properties but which do not require the use of a spatial variability model for the seismic action, the site average response spectrum shall be defined by combining, through a validated scientific method, the spectra corresponding to the differing soil conditions of the supports. (2) The site averaged response spectrum Sa may be defined as a weighted average of the appropriate site dependent response spectra and is determined by

(T)SR

R(T)S ii

jj

ia ∑ ∑

= (3.1)

where Ri is the reaction force on the base of pier i when the deck is subjected to a unit displacement while the base is kept immobile; Si is the site dependent response spectrum appropriate to the soil conditions at the foundation of pier i. Note: The average shall be computed separately for each of the two horizontal components and for the

vertical component.


(3) Alternatively the site averaged response spectrum may be substituted by an envelope spectrum obtained by considering, for each period, the highest value of the site dependent response spectra corresponding to the different soil conditions at the foundations of the bridge.

3.2.2.3 Site dependent power spectrum

(1)P The earthquake action can be described by a stochastic stationary gaussian process defined by a power spectrum and considered with a duration limited to a given time interval. This description of the motion shall be consistent with the site dependent response spectrum. Consistency between power spectrum and response spectrum shall be defined as equality between the response spectrum value and the mean value of the probability distribution of the largest extreme value (for the duration considered) of the response of a one degree of freedom oscillator with a corresponding natural frequency and viscous damping.

Note: The term extreme value refers to the absolute value of a maximum or a minimum value. It should be noted that in some cases (local) maximum values may have negative values and (local) minimum values may have positive values.

3.2.2.4 Time-history representation

(1)P When a non-linear time-history analysis is carried-out, at least three pairs of horizontal ground motion time-history components shall be used. The pairs should be selected from recorded events with magnitudes, source distances, and mechanisms consistent with those that define the design seismic action. (2) When the required number of pairs of appropriate recorded ground motions is not available, appropriate simulated accelerograms may be used in replacement of the missing recorded motions. (3)P Consistency to the applicable 5% damped design seismic spectrum shall be established by scaling the amplitude of motions as follows: • For each earthquake consisting of a pair of horizontal motions the SRSS spectrum shall

be created by taking the square root of the sum of squares of the 5%-damped spectra of each component.

• The spectrum of the ensemble of earthquakes shall be formed by taking the average value

of the SRSS spectra of the individual earthquakes of the previous step. • The ensemble spectrum shall be scaled so that it is not lower than 1,3 times the 5%-

damped design seismic spectrum, in the period range between 0,2T1 and 1,5 T1. Where T1 is the natural period of the fundamental mode of the structure in the case of a ductile


bridge, or the effective period (Teff) of the isolation system in the case of a bridge with seismic isolation (see 7.2).

• The scale factor derived in the previous step shall be applied to all individual seismic

motion components. (4)P The components of each pair of time histories shall be applied simultaneously. (5) When three component ground motion time-history recordings are used for non-linear time history analysis, scaling of the horizontal pairs of components may be carried out in accordance with (3)P above, independently from the scaling of the vertical components. The latter shall be effected so that the average of the relevant spectra of the ensemble is not lower by more than 10% of the 5% damped design vertical seismic spectrum, in the period range between 0,2Tv and 15Tv, where Tv is the period of the lowest mode where vertical motion of the structure prevails. (6) The use of pairs of horizontal ground motion recordings in combination with vertical recordings of different seismic motions, consistent with the requirements of (1)P above, is also allowed. The independent scaling of the pairs of horizontal recordings and of the vertical recordings shall be carried out as in (5) above.

3.2.2.5 Site dependent design spectrum for linear analysis

(1)P Both ductile and limited ductile structures shall be designed by performing linear analysis using a reduced response spectrum called the design spectrum as specified by 3.2.2.5 of EN 1998-1. 3.2.3 Six component model

3.2.3.1 General

(1)P The six component model of the earthquake motion at a point shall be developed from the probable contribution of the P, S, Rayleigh and Love waves to the total earthquake vibration.

Note: The simplified models referred to in Annex D may be used if geological discontinuities are not present.

3.2.3.2 Separation of the components of the seismic action

(1)P For the separation of the components of the seismic action the relevant provisions of 3.1.3 are applicable. However, the vertical component may, in general, be disregarded if the bridge is not particularly sensitive to vibrations in this direction; furthermore, the rotational components are usually not important and can also be disregarded.


3.3 Characterization of the spatial variability (1)P The spatial variability shall be considered when: • Geological discontinuities are present (e.g. soft soil contiguous to rock) • Marked topographical features are present; • The length of the bridge is greater than 600 m, even if there are no geological discontinuities

or marked topographical features.

Note: More detailed conditions and simplified models to take into account the spatial variability of the earthquake motion may be defined in the National Annex. Relevant information is presented in informative Annex D.

(2)P The spatial variability dealt with in this subclause concerns the continuous deformation of the ground, in the elastic or in the post-elastic range. However, in the case of strong earthquakes, discontinuous deformations, due to surface faulting or soil ruptures, may be induced. Measures to mitigate this hazard, such as the adoption of structural systems which minimize its effects, shall be taken. (See also 2.4 (10)).


4. ANALYSIS

4.1 Modelling 4.1.1 Dynamic degrees of freedom (1)P The model of the bridge and the selection of the dynamic degrees of freedom shall adequately represent the distribution of stiffness and mass so that all significant deformation modes and inertia forces are properly activated under the design seismic excitation. (2) It is sufficient, in most cases, to use two separate models in the analysis, one for modelling the behaviour in the longitudinal direction, and the other for the transverse direction. The cases when it is necessary to consider the vertical component of the seismic action are defined in 4.1.7. 4.1.2 Masses (1)P For the calculation of masses the mean values of the permanent masses and the quasi-permanent values of the masses corresponding to the variable actions shall be considered. (2) Distributed masses may be lumped at nodes according to the selected degrees of freedom. (3)P For design purposes the mean values of the permanent actions are identified by their characteristic values; the quasi -permanent values of variable actions are given by ψ2,1 Qk,1 where Qk,1 is the characteristic value of traffic load. In general and in accordance with EN1990-Annex A2 the value of ψ2,1=0 shall be used for bridges with normal traffic and foot bridges. (4)P For bridges with intense traffic the values of ψ2,1 may be set in the National Annex. These values shall be applied to the uniform load of Model 1 (LM 1) according to EN1991-2. Note: The following values are recommended.

• Road bridges ψ2,1 = 0,2 • Railway bridges ψ2,1 = 0,3

(5) When the piers are immersed in water, in the absence of a rigorous assessment of the hydrodynamic interaction, this effect may be estimated by taking into account an added mass of entrained water acting in the horizontal directions per unit length of the immersed pier, as described in Annex F. The hydrodynamic influence on the vertical seismic action may be omitted. 4.1.3 Element stiffness (1) For the estimation of element stiffnesses refer to 2.3.6.1. (2) Continuity slabs (see 2.3.2.1 (4)) should be included in the model of seismic analysis taking into account their excentricity relative to the deck axis and a reduced value of their


flexural stiffness. Unless this stiffness is estimated on the basis of the rotation of the relevant plastic hinges, a value of 25% of the uncracked flexural stiffness may be used. (3) For second order effects refer to 2.4 (5) and 5.4 (1). Significant second order effects may occur in bridges with slender piers and in special bridges like arch and cable - stayed bridges. 4.1.4 Modeling of the soil (1)P In general the supporting elements which transmit the seismic action from the soil to the deck shall be assumed as fixed to the foundation soil (see 3.2.1(2)). However soil-structure interaction effects may be considered according to EN1998-5, using appropriate impedances or appropriately defined soil springs. (2) Soil-structure interaction effects should be used when the displacement due to soil flexibility is greater than 30% of the total displacement at the centre of mass of the deck. (3) In those cases in which it is difficult to estimate reliable values for the mechanical properties of the soil, the analysis should be carried out using the estimated probable highest and lowest values. High estimates of soil stiffness should be used for computing the internal forces and low estimates for computing displacements. 4.1.5 Torsional effects (1)P Skew bridges (skew angle ϕ > 20o ) and bridges with a ratio B/L > 2,0 tend to rotate about the vertical axis, despite a theoretical coincidence of the centre of mass with the centre of stiffness. (L is the total length of the continuous deck and B is the width of the deck).

(2) Highly skewed bridges (ϕ > 45ο) should in general be avoided in high seismicity regions. If this is not possible, and the bridge is supported on the abutments through bearings, the actual horizontal stiffness of the bearings must be carefully modeled, taking into account the

L

B Figure 4.1

Skew bridge

φ


concentration of the vertical reactions near the obtuse angles. Alternatively an increased accidental eccentricity may be used. (3)P When using the Fundamental Mode Method (see 4.2.2) for the design of such bridges, the following equivalent static moment shall be considered to act about the vertical axis at the center of gravity of the deck: Mt = ± F e (4.1) where: F is the horizontal force determined according to the relation (4.14) e = ea + ed ea = 0,03L or 0,03B is the accidental mass eccentricity, and ed = 0,05L or 0,05B is an additional eccentricity reflecting the dynamic effect of simultaneous translational and torsional vibration; for the calculation of ea and ed the dimension L or B transverse to the direction of excitation shall be used. (4) When using a Full Dynamic Model (space model) the dynamic part of the torsional excitation is taken into account if the centre of mass is displaced by the accidental eccentricity ea in the most unfavorable direction and sense. However, the torsional effects may also be estimated using the static torsional moment of relation (4.1). (5)P The torsional resistance of a bridge structure shall not rely on the torsional rigidity of a single pier. In single span bridges the bearings shall be designed to resist the torsional effects. 4.1.6 Behaviour factors for linear analysis (1)P The standard procedure of the present code is an equivalent linear response spectrum analysis. The behaviour factor is defined globally for the entire structure and reflects its ductility capacity i.e. the capability of the ductile elements to withstand, with acceptable damage but without failure, seismic actions in the post-elastic range. The available levels of ductility are defined in 2.3.2. The capability of ductile elements to develop flexural plastic hinges is an essential requirement for the application of the values of the behaviour factor q defined in Table 4.1 for ductile behaviour.

Note: The linear analysis method, using an appropriate global force reduction factor (behaviour factor), is generally considered to be a reasonable compromise between the uncertainties intrinsic to the seismic problem and the relevant admissible errors on the one hand and the required effort for the analysis and design on the other.

(2) This capability is deemed to be ensured when the detailing rules of Section 6 are followed and capacity design according to 5.3 is performed. (3)P The maximum values of the behaviour factor q which may be used for the two horizontal seismic components are given in Table 4.1, depending on the post-elastic behaviour of the ductile elements, where the main energy dissipation takes place. If a bridge has various types of ductile elements, the q factor corresponding to the type-group having the major contribution to the seismic resistance shall be used. However, different q factors may be used in each horizontal direction.


Note: Use of behaviour factor values less than the maximum allowable given in Table 4.1 will normally lead to reduced ductility demands, which in general implies a reduction of potential damage. Such a use is therefore at the discretion of the Designer and the Owner .

Table 4.1: Maximum values of the behaviour factor q

Type of Ductile Elements

Seismic Behaviour

Limited Ductile

Ductile

Reinforced concrete piers: Vertical piers in bending (αs* ≥ 3,0) Inclined struts in bending Steel Piers: Vertical piers in bending Inclined struts in bending Piers with normal bracing Piers with eccentric bracing Abutments rigidly connected to the deck: In general Locked in structures (par. (9), (10)) Arches

1,5 1,2 1,5 1,2 1,5 - 1,5 1,0 1,2

3,5 λ(αs) 2,1 λ(αs) 3,5 2,0 2,5 3,5 1,5 1,0 2,0

*αs= L/h is the shear ratio of the pier, where L is the distance from the plastic hinge to the point of zero moment and h is the depth of the cross section in the direction of flexure of the plastic hinge. For αs ≥ 3 λ(αs) = 1,0

3 > αs ≥ 1,0 λ(αs) = 3sα

(4) For all bridges with regular seismic behaviour as defined in 4.1.8, when the detailing requirements given in Section 6 are met, the values of the q-factor given in Table 4.1 for Ductile Behaviour may be used without any special verification of the available ductility. When the requirements given in 6.5 are met, the values of the q-factor given in Table 4.1 for Limited Ductile Behaviour may be used without any special verification of the available ductility, regardless of the regularity or irregularity of the bridge. (5)P For reinforced concrete ductile elements the values of q-factors given in Table 4.1 are valid when the normalized axial force ηk defined in 5.3 (4) does not exceed 0,30. When 0,30 < ηk ≤ 0,60, even in a single ductile element, the value of the behaviour factor shall be reduced to:

1,01)(0,3

0,3kr ≥−

−−= qη

qq (4.2)

A value for qr = 1,0 (elastic behaviour) should be used for bridges in which the seismic force resisting system contains elements with ηk > 0,6.


(6)P The values of the q-factor for Ductile Behaviour given in Table 4.1 may be used only if the locations of all the relevant plastic hinges are accessible for inspection and repair. Otherwise, the values of Table 4.1 shall be multiplied by 0,6; however final q-values less than 1,0 need not be used.

Note: The term “accessible”, as used in the paragraph above, has the meaning of “accessible even with reasonable difficulty”. The foot of a pier shaft located in backfill, even at substantial depth, is considered to be “accessible”. On the contrary, the foot of pier shaft immersed in deep water, or the heads of piles beneath an extensive pile cap, should not be considered “accessible”.

(7) When energy dissipation is intended to occur at plastic hinges located in piles, which are designed for ductile behaviour, and at points which are not accessible, a final q-value of 2,1 shall be used for vertical piles and 1,5 for inclined piles (see also 5.4.2(5) of EN1998-5). (8) Regarding plastic hinge formation in the deck, see 2.3.2.1 (4).

Note: No plastic hinges will in general develop in piers flexibly connected to the deck, in the direction under consideration. A similar situation will occur in individual piers having very low stiffness in comparison to the other piers (see 2.3.2.1 (6) and (7)). Such elements have negligible contribution in resisting the seismic actions, and therefore do not affect the q-factor (see 4.1.6 (3)P).

(9) Bridge structures whose mass follows essentially the horizontal seismic motion of the ground (“locked-in” structures), do not experience significant amplification of the horizontal ground acceleration. Such structures are characterized by a very low value of the natural period in the horizontal directions (T ≤ 0,03 s). The inertial response of these structures in the horizontal directions may be assessed using the design value of the seismic ground acceleration and q = 1. Abutments flexibly connected to the deck belong to this category. (10) Bridge structures consisting of an essentially horizontal deck, rigidly connected to both abutments (either monolithically or through fixed bearings or links), may be considered to belong to the category of the previous paragraph (9) (without need to check the natural period) if the abutments are laterally encased, at least over 80 % of their area, in stiff natural soil formations. If above conditions are not met, then the soil interaction at the abutments should be included in the model, using realistic soil stiffness parameters. In case T > 0,03 s, then the normal acceleration response spectrum with q = 1,50 should be used. (11) When the main part of the design seismic action is resisted by elastomeric bearings the flexibility of the bearings imposes a practically elastic behaviour of the system. Such bridges shall be designed according to Section 7.

Note: The potential formation of plastic hinges in secondary deck elements (continuity slabs) is allowed but should not be considered to increase the value of q.

(12)P The behaviour factor for the analysis in the vertical direction shall always be taken equal to 1,0.


4.1.7 Vertical component of the seismic action (1) The effects of the vertical seismic component on the piers may, as a rule, be omitted in zones of low and moderate seismicity. In zones of high seismicity these effects need only be investigated in the exceptional cases when the piers are subjected to high bending stresses due to the permanent actions of the deck, or when the bridge is located within 5km from an active seismotectonic fault. (2)P The effects of the vertical seismic component in the upward direction in prestressed concrete decks, shall be always investigated. (3)P The effects of the vertical seismic component on bearings and links shall be assessed in all cases. (4) The estimation of the effects of the vertical component may be carried out using the Fundamental Mode Method and the Flexible Deck Model (see 4.2.2.4). 4.1.8 Regular and irregular seismic behaviour of ductile bridges (1) Designating by MEd,i the maximum value of design moment under the seismic load combinations at the intended location of plastic hinge of ductile member i, and by MRd,i the design flexural resistance of the same section, with its actual reinforcement, under the concurrent action of the other action effects of the seismic load combination, then the required local force reduction factor ri associated with member i , under the specific seismic action is:

iRd,

iEd,i M

Mqr = (4.3)

Note: Since MEdi ≤ MRdi , it follows that ri

≤ q

(2) A bridge shall be considered to have a regular seismic behaviour, in the direction under consideration, when following condition is satisfied

omin

max ρrr

ρ ≤= (4.4)

where rmin = minimum (ri) and rmax = maximum (ri) for all ductile elements i and ρo = 2,00. One or more ductile elements (piers) may be exempted from the above calculation of rmin and rmax, if the sum of their shear contributions does not exceed 20% of the total seismic shear in the direction under consideration.


Note: When in a regular bridge rmax is substantially lower than q, the design does not fully exploit the allowable maximum q-values. When rmax = 1,0 the bridge responds elastically to the design earthquake considered.

(3) Bridges not meeting condition (4.4), should be considered to have irregular seismic behaviour, in the direction under consideration. Such bridges should either be checked to satisfy the requirements of Annex H or should be designed using a reduced q-value:

1,0≥=o

r ρρqq (4.5)

4.2 Methods of analysis 4.2.1 Linear dynamic analysis - Response spectrum method

4.2.1.1 Definition, field of application

(1) The Response Spectrum Analysis is an elastic analysis of the peak dynamic responses of all significant modes of the structure, using the ordinates of the site-dependent design spectrum (see 3.2.2.5). The overall response is obtained by statistical combination of the maximum modal contributions. Such an analysis may be applied in all cases in which a linear analysis is allowed. (2)P The earthquake action effects shall be determined from an appropriate discrete linear model (Full Dynamic Model), idealized in accordance with the laws of mechanics and the principles of structural analysis, and compatible with an adequate idealization of the earthquake action.

4.2.1.2 Significant modes

(1)P All modes having significant contribution to the total structural response shall be considered. (2) For bridges in which the total mass M can be considered as a sum of "effective modal masses" Mi, the above criterion is considered to be satisfied if the sum of the effective modal masses, for the modes considered (ΣMi)c, amounts at least to 90% of the total mass of the bridge. (3) In cases that the above condition is not satisfied after consideration of all modes with T ≥ 0,033s, the number of modes considered may be deemed acceptable provided that: • (ΣMi)c/M ≥ 0,70 • The final values of the seismic effects are multiplied by M/(ΣMi)c


4.2.1.3 Combination of modal responses

(1)P The probable maximum value E of an action effect (force, displacement etc.), shall be taken in general equal to the square root of the sum of squares of the modal responses Ei (SRSS-rule) 2

iΣΕE = (4.6) This action effect shall be assumed to act in both senses. (2)P When two modes have closely spaced natural periods Tj ≤ Ti, with the ratio ρ = Tj/Ti exceeding the value 0,1 /(0,1 + ξ'), where ξ' is the viscous damping ratio (see (3) below), the SRSS rule becomes unconservative and more accurate rules shall be applied. (3) For the above case the method of the Complete Quadratic Combination (CQC) may be used:

jijijiΕrΕΣΣΕ = (4.7)

with i = 1 ... n , j = 1 ... n with the correlation factor

( ) ( )2'222

3/2'2

j 14+1

+(18

ρ+ρξρ-

ρ)ρξ=ri (4.8)

where: ρ = Tj/Τi and ξ is the viscous damping ratio (4) When the differential displacement along the base of the bridge can induce substantial stresses in the structure, the value of the earthquake action effects can be determined in the case of application of the SRSS-rule as 2

mmm2ii )d(kΣΕΣΕ += (4.9)

and in the case of application of the CQC-method as

2

mmmjijiji )d(kΣΕrΕΣΣΕ += (4.10) where km is the effect of the m-th independent motion and dm is the assymtotic value of the spectrum for the m-th motion for large periods, expressed in displacements.


4.2.1.4 Combination of the components of seismic action

(1) The probable maximum action effect E, due to the simultaneous occurrence of seismic actions along the horizontal axes X, Y and the vertical axis Z, may be estimated from the maximum action effects Ex, Ey and Ez due to independent seismic action along each axis, as follows: EE+EE 2

z2y

2x += (4.11)

(2) Alternatively it is sufficient to use as design seismic action AEd the most adverse of the following combinations: AEx "+" 0,30AEy "+" 0,30 AEz 0,30AEx "+" AEy "+" 0,30AEz 0,30AEx "+" 0,30AEy "+" AEz (4.12) where AEx, AEy and AEz are the seismic actions in each direction X, Y and Z respectively. AEz should be considered according to the requirements of 4.1.7. 4.2.2 Fundamental mode method

4.2.2.1 Definition

(1) Equivalent static seismic forces are derived from the inertia forces corresponding to the fundamental mode and natural period of the structure in the direction under consideration, using the relevant ordinate of the site dependant design spectrum. The method includes also simplifications regarding the shape of the first mode and the estimation of the fundamental period. (2) Depending on the particular characteristics of the bridge, this method may be applied using three different approaches for the model, namely: • the Rigid Deck Model • the Flexible Deck Model • the Individual Pier Model (3)P The rules of 4.2.1.4 for the combination of the components of seismic action shall be applied.

4.2.2.2 Field of application

(1) The method may be applied in all cases in which the dynamic behaviour of the structure can be sufficiently approximated by a single dynamic degree of freedom model. This condition is considered to be satisfied in the following cases:


(a) In the longitudinal direction of approximately straight bridges with continuous deck, when the seismic forces are carried by piers whose total mass is less than 1/5 of the mass of the deck.

(b) In the transverse direction of case (a) when the structural system is approximately

symmetrical about the centre of the deck, i.e. when the theoretical eccentricity eo between the centre of stiffness of the supporting elements and the centre of mass of the deck does not exceed 5% of the length of the deck (L).

(c) In the case of piers carrying simply supported spans when no significant interaction between

piers is expected and the total mass of each pier is less than 1/5 of the mass of the part of the deck carried by the pier.

4.2.2.3 Rigid deck model

(1) This model may be applied only when - under the earthquake action - the deformation of the deck in a horizontal plane is negligible compared to the displacements of the pier tops. This is always valid in the longitudinal direction of approximately straight bridges with continuous deck. In the transverse direction the deck may be assumed rigid if L/B ≤ 4,0 or, in general, if the following condition is satisfied:

0,20a

d ≤d∆

(4.13)

where: L is the total length of the continuous deck B is the width of the deck and ∆d and da are respectively the maximum difference and the average of the displacements in the transverse direction of all pier tops under the transverse seismic action or under the action of a transverse load of similar distibution. (2)P The earthquake effects shall be determined by applying at the deck a horizontal equivalent static force F given by the expression:

g

(T)GSF d= (4.14)

where: G is the total effective weight of the structure, equal to the weight of the deck plus the

weight of the upper half of the piers. Sd(T) is the spectral acceleration of the design spectrum (3.2.2.5) corresponding to the

fundamental period T of the bridge, estimated as:


gKGπT 2= (4.15)

where Κ=ΣΚi is the stiffness of the system, equal to the sum of the stiffnesses of the resisting elements. (3) In the transverse direction the force F may be distributed along the deck proportionally to the distribution of the effective masses.

4.2.2.4 Flexible deck model

(1)P The Flexible Deck Model shall be used when the condition (4.12) is not satisfied. (2) In the absence of a more rigorous calculation, the fundamental period of the structure, in the direction under examination, may be estimated by the Rayleigh method, using a generalized single degree of freedom system, as follows:

∑

∑=

ii

i22i

dGg

dGπT (4.16)

where: Gi is the weight concentrated at the i-th nodal point di is the displacement in the direction under examination when the structure is acted upon by

forces Gi acting at all nodal points in the direction under consideration. (3)P The earthquake effects shall be determined by applying at all nodal points horizontal forces Fi given by :

iid

2

2

i Gdg(T)S

gT4πF = (4.17)

where : T is the period of the fundamental mode of vibration for the direction under consideration, Gi is the weight concentrated at the i-th point, di is the displacement of the i-th nodal point, approximating the shape of the first mode (it

may be taken equal to the values determined in (2) above), Sd(T) is the spectral acceleration of the design spectrum (3.2.2.5), and g is the acceleration of gravity.

4.2.2.5 Torsional effects in the transverse direction (rotation about the vertical axis)

(1) When the Rigid or the Flexible Deck Model is used in the transverse direction of a bridge, torsional effects may be estimated by applying a static torsional moment Mt according to relation (4.1) of 4.1.5 (3). The relevant eccentricity shall be estimated as follows:


e = eo + ea (4.18) where: eo is the theoretical eccentricity (see case (b) of 4.2.2.2) ea = 0,05L is an accidental eccentricity accounting for accidental and dynamic amplification effects (2) The force F may be determined either according to equation (4.14 or as ΣFi according to equation (4.17). The moment Mt may be distributed to the supporting elements using the Rigid Deck Model.

4.2.2.6 Individual pier model

(1) In some cases, earthquake action in the transverse direction of bridges is resisted mainly by the piers, and no strong interaction between adjacent piers develops. In such cases the earthquake effects acting on the i-th pier may be approximated by considering the action of an equivalent static force.

g

)(TSGF idi

i = (4.19)

where Gi is the effective weight attributed to pier i and

gK

GπTi

i2i = (4.20)

is the fundamental period of the same pier. (2) This simplification may be applied as a first approximation for preliminary analyses, when the following condition is met for all adjacent piers i and i+1 0,90 ≤ Ti/Ti+1 ≤ 1,10 (4.21) Otherwise a redistribution of the effective masses attributed to each pier, leading to the satisfaction of the above condition, is required. 4.2.3 Alternative linear methods

4.2.3.1 Power spectrum analysis

(1)P A linear stochastic analysis of the structure shall be performed, either by applying modal analysis or frequency dependent response matrices, using as input the acceleration power density spectrum (see 3.2.2.3 ).


(2)P The elastic action effects shall be defined as the mean value of the probability distribution of the largest extreme value of the response for the duration considered in the earthquake model. (3)P The design values shall be determined by dividing the elastic effects by the appropriate behaviour factor q, and ductile response shall be assured by compliance with the relevant sections of this Code. (4) The method has the same field of application as the Response Spectrum Analysis.

4.2.3.2 Time series analysis

(1)P In a time series analysis the effects of the earthquake action shall be identified with the sample average of the extreme response computed for each accelerogram in the sample. For the definition of time histories see 3.2.2.4. 4.2.4 Non - linear dynamic time-history analysis

4.2.4.1 General

(1)P The structure´s time dependent response shall be obtained through direct numerical integration of its non-linear differential equations of motion. The seismic input shall consist of ground motion time histories (accelerograms, see 3.2.2.4). The effects of gravity loads and other concurrent quasi-permanent actions, as well as second order effects shall be taken into account. (2)P Unless otherwise specified in this Part, this method can be used only in combination with a standard response spectrum analysis to provide insight into the post - elastic response and comparison between required and available local ductilities. Generally, the results of the non-linear analysis shall not be used to relax requirements resulting from the response spectrum analysis. However, in the case of bridges with isolating devices (Section 7) and irregular bridges (4.1.8) lower results from a rigorous time-history analysis may be substituted for the results of the response spectrum analysis.

4.2.4.2 Ground motions and design combination

(1)P The provisions of 3.2.2.4 are applicable. (2)P The provisions of 5.5(1) and 4.1.2 are applicable.

4.2.4.3 Design response effects

(1)P When non-linear dynamic analysis is performed for at least 7 independent pairs of horizontal ground motions, the average of the individual responses may be assumed as design response effects, except if otherwise required in this part. When less than 7 input motions are


analyzed, the maximum responses of the ensemble should be assumed as design response effects.

4.2.4.4 Ductile structures

(1) Objectives: The main objectives of a non-linear time history analysis of a ductile bridge are the following: • identification of the actually required pattern of plastic hinge formation • estimation and verification of the probable demands for post-yield deformation of the plastic

hinges, and estimation of the displacement demands • strength verification against non-ductile failure modes of the members and against soil

failure (2) Requirements: For a ductile structure subjected locally to substantial excursions in post-yield deformations, that are necessary for achieving reduction of the seismic action, pursuing above objectives requires the following:

(a) A realistic assessment of the extent of the structure whose response remains below yield. Such an assessment should be based on probable values of the yield stresses and strains of the materials. Guidance for the selection of these values is given in Annex E.

(b) In the regions of plastic hinges, the stress-strain diagrams for both concrete and

reinforcement or structural steel, should reflect the probable post-yield behaviour, taking into account confinement, when relevant, for concrete, and strain hardening and/or local buckling effects for steel (see Annex E). The hysteretic loop shape should be properly modeled, taking into account strength and stiffness degradation and hysteretic pinching if indicated by appropriate laboratory tests.

(c) The verification that deformation demands are safely lower than the capacities of the plastic

hinges, should be performed in terms of plastic hinge rotation demands θp,E, by comparison to relevant design rotation capacities θp,d, as follows:

θp,E ≤ θp,d (4.22) Design plastic rotation capacities θp,d, should be derived from relevant test results or

calculated from ultimate curvatures, by deviding the probable value θp,u by factor γR,p, reflecting local defects of the structure, uncertainties of the model and/or the dispersion of relevant test results, as follows:

pR,

up,dp, γ

θθ = (4.23)

Guidance for the estimation of θp,d is given in Annex E. The same condition should be

checked for other deformation demands and capacities of dissipative zones of steel


structures (elongation of tensile elements in diagonals and shear deformation of shear panels in eccentric bracings).

(d) Member strength verification against bending with axial force is not needed, as such a

verification is inherent in the non-linear analysis procedure, according to point (a) above. However it shall be verified that no significant yield occurs in the deck (see H.2.4).

(e) Verification of all members against non-ductile failure modes (shear of members and shear

of joints adjacent to plastic hinges) as well as of foundation failure, should be performed, according to the relevant rules of Section 5 of this part, assuming as design actions (in lieu of the capacity design effects), the maximum values of the responses of the ensemble of the analyses for the ground motions used (maxAEd). These values should not exceed the design resistances Rd (= Rk/γM) of the corresponding sections

maxAE,d ≤ Rd (4.24)

4.2.4.5 Bridges with seismic isolation

(1) The objective of the analysis in this case is the realistic assessment of the displacement and force demands: • taking into proper consideration the variability of the properties of the isolators, and • ensuring that the isolated structure remains essentially elastic (2) The provisions of Section 7 are applicable. 4.2.5 Static non-linear analysis (pushover analysis) (1)P Pushover analysis is a static non-linear analysis of the structure under constant vertical (gravity) loads and monotonically increased horizontal loads, representing the effect of an horizontal seismic component. Second order effects shall be accounted for. The horizontal loads are increased until a target displacement is reached at a reference point. (2) The main objectives of the analysis are the following: • estimation of the sequence and the final patern of plastic hinge formation • estimation of the redistribution of forces following the formation of plastic hinges • assessment of the force-displacement curve of the structure and of the deformation demands

of the plastic hinges up to the target displacement (3) The method may be applied to the entire bridge structure or to individual components. (4) The requirements of 4.2.4.4 (2) are applicable, with the exception of the requirement for modeling of the hysteretic loop shape mentioned in (b). (5) Guidance for application of this method is given in H.2 of Annex H.



5. STRENGTH VERIFICATION

5.1 Scope (1)P The provisions of this Section are applicable to the earthquake resisting system of bridges designed by an equivalent linear method taking into account a ductile or limited ductile behaviour of the structure (see 4.1.5). For bridges provided with isolating devices, Section 7 shall be applied. For verifications on the basis of results of non-linear analysis, provisions 4.2.4 are applicable. 5.2 Materials and Design strength 5.2.1 Materials

(1)P In bridges of ductile behaviour with q > 1,50, concrete elements where plastic hinges may occur, shall be reinforced using steel of Class C of Table 3.3 of EN1992-1 (2) Concrete elements of bridges of ductile behaviour, where no plastic hinge may occur (as a consequence of capacity design), as well as all concrete elements of bridges of limited ductile behaviour (q ≤ 1,50) or of bridges with seismic isolation according to Section 7, may be reinforced using steel of Class B of Table 3.3 of of EN 1992-1 (3)P Structural steel elements for all bridges shall conform to 6.2 of EN 1998-1 5.2.2 Design strength

(1)P The material safety factors γM as defined in 1992-2, 1993-2 and 1994-2 for the fundamental load combinations shall also be used for strength verifications under seismic load combinations and capacity design effects. 5.3 Capacity design (1)P For structures of ductile behaviour, capacity design effects (FC) shall be calculated by analysing the intended plastic mechanism under the permanent actions and the level of seismic action at which all intended flexural hinges have developed bending moments equal to an appropriate upper fractile of their flexural resistance, called the overstrength moment Mo. (2) The capacity design effects need not be taken greater than these resulting from the design seismic combination (cl. 5.5) where the design effects AEd are multiplied ny the q factor used. (3)P The overstrength moment of a section shall be calculated as: Mo = γoΜRd (5.1) where: γo is the overstrength factor MRd is the design flexural strength of the section, in the selected direction and sense, based on

the actual section geometry and reinforcement configuration and quantity (with γMvalues for fundamental load combinations). In determining MRd, biaxial bending under the


permanent effects, and the seismic effects corresponding to the design seismic action in the selected direction and sense, shall be considered.

(4)P The value of the overstrength factor shall be taken in general as: γo = 1,35 (5.2) In the case of reinforced concrete sections, with special confining reinforcement according to 6.2.1, in which the value of the normalized axial force ηk = NEd/(Acfck) (5.3) exceeds 0,1, the value of the overstrength factor shall be increased to: γo= 1,35 [1+2(ηk-0,1)2] (5.4) where: NEd is the value of the axial force at the plastic hinge corresponding to the design seismic

combination, positive if compressive Ac is the area of the section and fck is the characteristic concrete strength. (5)P Within members containing plastic hinge(s), the capacity design bending moment Mo at the vicinity of the hinge (see Figure 5.1) shall not be assumed greater than the relevant design flexural resistance MRd of the hinge assessed according to 5.6.3.1.

Note 1: The MRd-curves shown in Fig. 5.1 correspond to a pier with variable cross section (increasing downwards). In case of constant cross section, since the reinforcement is constant, MRd is also constant. Note 2: For Lh see 6.2.1.4.

γ0MRd t

MRd b γ0MRd b ME b

MRd t ME t Deck

Pier

P.H

P.H

P.H: Plastic hinge

Figure 5.1 Capacity design moments within member containing plastic hinges

Lh


(6)P Capacity design effects shall be calculated in general for each sense of the seismic action in both the longitudinal and the transverse directions. A relevant procedure and simplifications are given in Annex G. (7)P When sliding bearings participate in the plastic mechanism, their capacity shall be assumed equal to γofRdf ,where γof = 1,30 is a magnification factor for friction due to ageing effects and Rdf is the maximum design friction force of the bearing. (8)P In bridges intended to have ductile behavior, in the case of members where no plastic hinges are intended to form and which resist shear forces from elastomeric bearings, the capacity design effects shall be calculated on the basis of the maximum deformation of the elastomeric bearings corresponding to the design displacement of the deck. A 30% increase of the bearing stiffness shall be used in these cases. 5.4 Second order effects (1) For linear analysis, approximate methods may be used for estimating the influence of second order effects on the critical sections (plastic hinges). Such methods should be reasonably conservative taking into account the cyclic character of the seismic action.

Note: Approximate methods for estimating second order effects under seismic actions may be defined in the National Annex. In the absence of a more accurate procedure, the increase of bending moments of the plastic hinge section due to second order effects, may be assumed equal to:

EdEd21 Ndq∆Μ +

= (5.5)

where NEd is the axial force and dEd is the relative transverse displacement of the ends of the ductile member under consideration.

5.5 Design seismic combination (1)P The design value of action effects Ed, in the seismic design situation shall be derived from the following combination of actions: Gk "+" Pk "+" AEd "+" ψ21Q1k "+" Q2 (5.6) where: Gk are the permanent loads with their characteristic values, Pk is the characteristic value of prestressing after all losses, AEd is the most unfavorable combination of the components of the earthquake action according

to 4.2.1.4, Q1k is the characteristic value of the traffic load, and ψ21 is the combination factor according to 4.1.2 (3). Q2 is the quasi permanent value of actions of long duration (e.g. earth pressure, buoyancy,

currents etc.)


Note: Actions of long duration are considered to be concurrent with the design earthquake. (2)P Seismic action effects need not be combined with action effects due to imposed deformations (temperature variation, shrinkage, settlements of supports, ground residual movements due to seismic faulting). (3)P An exception to the rule stated above is the case of bridges in which the seismic action is resisted by elastomeric laminated bearings (see also 6.6.2.3(3)). In this case elastic behaviour of the system must be assumed and the action effects due to imposed deformations must be accounted for.

Note: It is noted that in the above case the displacement due to the creep does not normally induce additional stresses to the system and can therefore be omitted. Creep also reduces the effective value of long-term imposed deformations (e.g. shrinkage), which induces stresses in the structure .

(4)P For wind and snow actions the value ψ21=0 shall be assumed. 5.6 Resistance verification of concrete sections

5.6.1 Design effects (1) When the resistance of a section depends significantly on the interaction of more than one action effect (e.g. bending moment and axial force) it is sufficient that the Ultimate Limit State conditions, given in 5.6.2 and 5.6.3, are satisfied separately by the extreme (max. or min.) value of each action, taking into account the interaction with the coincidental accompanying values of the other actions. 5.6.2 Structures of limited ductile behaviour (1)P Ed ≤ Rd (5.7) where: Ed is the design action effect under the seismic load combination including second order

effects, and Rd is the design resistance of the section. (2)P In regions of moderate to high seismicity the shear resistance shall be verified using Ed from (5.6) with the seismic action effects (AEd) multiplied by the value q used in the analysis. Note: For regions of low seismicity see Note in 2.3.7. 5.6.3 Structures of ductile behaviour

5.6.3.1 Flexural resistance of sections of plastic hinges

(1)P MEd ≤ MRd (5.8)


where: MEd is the design moment under the seismic load combination, including second order effects,

and MRd is the design flexural resistance of the section, as defined in cl. 5.3(3). (2)P The longitudinal reinforcement of the member containing the hinge shall remain constant and fully effective over the length lh indicated in Figure 5.1 (see also 6.2.1.4).

5.6.3.2 Flexural resistance of sections outside the region of plastic hinges

(1)P MC ≤ MRd (5.9) where : MC is the capacity design moment as defined in 5.3, and MRd is the design resistance of the section, taking into account the interaction of the

corresponding design effects (axial force and when applicable the bending moment in the other direction).

Note: As a consequence of 5.3 (5)P, the cross section and the longitudinal reinforcement of the plastic hinge setion shall not be affected by the capacity design verification (Condition (5.9)).

5.6.3.3 Shear resistance of elements outside the region of plastic hinges

(1)P The shear verifications shall be carried out according to the provisions of 6.2 of EN 1992-1, assuming as design action effects the capacity design effects defined by 5.3. (2) For circular concrete sections of radius r where the longitudinal reinforcement is distributed over a circle with radius rs, and in the absence of a rigorous assessment, the effective depth:

πr

rd se

2+= (5.10)

may be used instead of d in the relevant expressions for the shear resistance. The value of inner lever arm z may be assumed as z = 0,9de.

5.6.3.4 Shear resistance of plastic hinges

(1)P Paragraph (1)P of 5.6.3.3 is applicable. (2)P The angle θ between concrete compression strut and the main tension chord shall be assumed equal to 45o. (3)P The dimensions of the confined concrete shall be used in lieu of the section dimensions bw and d.


(4) Paragraph 5.6.3.3 (2) may be applied on the dimensions of the confined core. (5) For elements with shear ratio αs < 2,0 (see Table 4.1 for definition of αs), verifications of the pier against diagonal tension and sliding failure should be carried out in accordance with 5.5.3.4.3 and 5.5.3.4.4 of EN 1998-1 respectively, assuming as design actions the capacity design effects.

5.6.3.5 Verification of joints adjacent to plastic hinges

a. General (1)P Joints between vertical ductile piers and deck or foundation elements, adjacent to a plastic hinge, shall be designed in shear to resist the capacity design effects of the plastic joint in the relevant direction. The element(s) framing to the pier, are referred to in the following paragraphs as “beam(s)”, while the pier is refered to as “column”. (2)P For a vertical solid column of width bc, transverse to the direction of flexure of the plastic hinge, and of depth hc, the design width of the joint shall be assumed as follows: • when the column frames into a slab or a transverse rib of a hollow slab bj = bc + 0,5hc (5.11) • when the column frames directly on a longitudinal web of width bw (bw is parallel to bc) bj = min(bw,bc + 0,5hc) (5.12) • for circular column of diameter dc, above definitions are applicable assuming bc = hc = 0,9dc b. Joint forces and stresses

γ0TRc

x

z y

γ0MRd

zc

zb

Forces on the joint Vjz

Vjx Njx

Njz

Internal forces

zc

hc

hb

Figure 5.2 : Joint forces

Vb1C Vb2C

Plastic hinge γ0CRc


(1)P The design vertical shear of the joint, Vjz, shall be assumed as: Vjz = γoTRc – Vb1C (5.13) where TRc is the resultant force of the tensile reinforcement of the column corresponding to the

design flexural resistance MRd of the plastic hinge according to 5.3 (3)P, and γ0 is the overstrength factor according to 5.3 (3)P and (4)P (capacity design), and V1bC is the shear force of the beam adjacent to the tensile face of the column, corresponding to the capacity design effects of the plastic hinge

(2) The design horizontal shear of the joint Vjx may be calculated from the relation (see Fig. 5.2) Vjx zb = Vjz zc (5.14) Where zc and zb are the internal force lever arms of the plastic hinge and the beam end sections, respectively. Unless a rigorous assessment is performed, 0,9 of the relevant effective section depths may be assumed (see 5.6.3.3 and 5.6.3.4). (2) The shear verification should be carried out at the center of the joint, where, in addition to Vjz and Vjx, the influence of following axial forces may be taken into account:

• vertical axial joint force Njz equal to:

cGj

cjz 2

Nb

bN = (5.15)

where NcG is the axial force of the column under permanent actions • horizontal force Njx equal to the capacity design axial force effects in the beam, including the

effects of longitudinal prestressing after all losses, if such axial forces are actually effective throughout the width bj of the joint.

• horizontal force Njy in the transverse direction equal to the effect of transverse prestressing

after all losses, effective within the depth hc, if such prestressing is provided. (4) For the joint verification following average nominal stresses are derived.

Shear stresses: vj = vx = vz = bj

jz

cj

jx

zbV

zbV

= (5.16)

Axial stresses: cj

jzz hb

Nn = (5.17)

bj

jxx hb

Nn = (5.18)


c

jyy

bhhN

n = (5.19)

Note: As pointed out in 5.3 (6), the capacity design, and therefore the relevant joint verification, should be carried out in both senses of the deismic action. It is also noted that at knee-joints (e.g. over the end column of a multicolumn bent in the transverse bridge direction), the sense of MRd and Vb1C may be opposite to that shown in Figure 5.2 and Njx may be tensile.

c. Verifications (1) If the average shear stress in the joint, vj, does not exceed the cracking shear capacity of the joint, vj,cr, as indicated by the subsequent condition (5.20), then minimum reinforcement should be provided, according to subsequent paragraph (6).

vj ≤ vj,cr = ≤ 1,50fcdt (5.20)

where fctd = fctk0,05/γc is the design tensile strength of concrete. (2)P The diagonal compression induced in the joint by the diagonal strut mechanism shall not exceed the compressive strength of concrete in the presence of transverse tensile strains taking into account also confining pressures and reinforcement. (3) In the absence of a more precise model, the requirement of (2)P above may be satisfied by means of the subsequent condition: vj ≤ vj,Rd = 0,5αcvfcd (5.21) where, v = 0,6 (1-(fck/250)) (with fck in MN/m2) (5.22) αc accounts for the effects of, potentially existing, confining pressure (njy) and/or reinforcement (ρy) in the transverse direction y, on the compressive strength of the diagonal strut: αc = 1 + 2(njy + ρyfsd)/fcd ≤ 1,5 (5.23) ρy = Asy/(hchb) is the reinforcement ratio of closed stirrups potentially provided in the transverse direction of the joint panel (orthogonal to the plane of action), and fsd = 300 MN/m2 is a reduced stress of this reinforcement, for reasons of limitation of cracking. (4) Adequate reinforcement, both horizontal and vertical, shall be provided in the joint, to carry the design shear force. In the absence of a more precise model, this requirement may be satisfied by providing horizontal and vertical reinforcement amounts, ρx and ρz, respectively, such that:

)fn)

fnf

ctd

z

ctd

xctd ++ (1(1


sy

xjx f

n-vρ = (5.24)

sy

zjz f

n-vρ = (5.25)

where,

bj

sxx hb

Aρ = is the reinforcement ratio in the joint panel in horizontal direction,

cj

szz hb

Aρ = is the reinforcement ratio in the joint panel in vertical direction, and

fsy, is the design yield strength of the joint reinforcement. (5)P The shear reinforcement ratio ρx or ρy shall not exceed the maximum value:

sy

max 2cd

fνf

ρ = (5.26)

(6)P A minimum amount of shear reinforcement shall be provided in the joint panel, in both the horizontal and transverse directions, in the form of closed stirrups. The required minimum amount is,

sy

ctdmin f

fρ = (5.27)

d. Reinforcement arrangement (1) Vertical stirrups should enclose the longitudinal beam reinforcement at the face opposite to the column. Horizontal stirrups should enclose the column vertical reinforcement, as well as beam horizontal bars anchored into the joint. Continuation of column stirrups/hoops into the joint is recommended. (2) Up to 50% of the total amount of vertical stirrups required in the joint may be U-bars, enclosing the longitudinal beam reinforcement at the face opposite to the column (see Fig. 5.3).

(3) 50% of the bars of the top and bottom bending reinforcement of the beams, when continuous through the joint body and adequately anchored beyond it, may be taken into asccount for covering the required horizontal joint reinforcement area Asx. (4) The longitudinal (vertical) column reinforcement should reach as far as possible into the beam, ending just before the reinforcement layers of the beam at the face opposite to the column-beam interface. In the direction of flexure of the plastic hinge, the bars of both tensile regions of the column should be anchored by a rectangular hook directed towards the center of the column.


(5) When the amount of required reinforcement Asz and/or Asx, according to relations (5.25) and (5.24) is so high as to endanger the constructibility of the joint, then the alternative arrangement, described in the following paragraphs (5) and (6), may be applied (see Fig. 5.3). (6) Vertical stirrups of amount ρ1z ≥ ρmin, acceptable from the constructibility point of view, should be placed within the joint body. The remaining area Aszb = (ρz - ρ1z)bjhb, should be placed on each side of the beam, within the joint width bj and not further than 0,5hb from the corresponding column face. (7) The horizontal stirrups, placed within the joint body, may be reduced by ∆Asx, and the tensile reinforcement requirements of the beam fibers at the extension of the beam-column interface, should be increased by: ∆Asx = 0,5 ρjz bj hb (5.28) in addition to the reinforcement required in the relevant sections, by the verification in flexure under capacity design effects. Additional bars, to cover this requirement, should be placed within the joint width bj, and be adequately anchored, so as to be fully effective at a distance hb from the column face.

Stirrups in common areas count in both directions hb/2 hb/2

hb

hc

≤ 50%

Beam-column interface

bj

Aszb

Areas for Aszb hb/2

hb/2

hb/2 hb/2

Plan view Plastic hinge in x-direction

Plan view Plastic hinge in x- and y-directions

Figure 5.3 Alternative arrangement of reinforcement

y

x

∆Asx


5.7 Resistance verification for steel and composite elements 5.7.1 Steel piers

5.7.1.1 General

(1)P Dissipative zones are allowed only in the piers and not in the deck. (2)P The provisions of 6.5.2, 6.5.4 and 6.5.5 of EN 1998-1 are applicable for dissipative structures. (3) The provisions of 6.5.3 of EN 1998-1 are applicable. However cross-sectional class 3 is allowed only when q ≤ 1,5. (4) The provisions of 6.9 of EN 1998-1 are applicable for all bridge piers.

5.7.1.2 Piers as moment resisting frames

(1)P In dissipative frames the design values of the axial force NEd and shear forces VE,d shall be assumed equal to the capacity design actions NC and VC respectively as the latter are defined in 5.3. (2)P The design of the sections of plastic hinges both in beams and columns of the pier shall satisfy the conditions (6.3), (6.4) and (6.5) of 6.6.2 (2) and (6.8) of 6.6.3 (6) of EN 1998-1, using the values of NEd and VEd as defined in the previous paragraph.

5.7.1.3 Piers as frames with concentric bracings

(1)P The provisions of 6.7 of En 1998-1 are applicable with the following modifications: • The design values for the axial force and shear shall be as per 5.3 considering as force of all

diagonals the overstrength γoNpl,d of the corresponding to the weakest diagonal (see 5.3 for γo).

• The second part of condition (6.10) of (6.7.4 of En 1998-1 shall be replaced by the capacity

design action NEd = NC

5.7.1.4 Piers as frames with eccentric bracings

(1)P The provisions of 6.8 of EN 1998-1 are applicable. 5.7.2 Steel or composite deck (1)P For dissipative piers (q > 1,5) the deck shall be verified assuming the action of the design seismic combination ASd equal to the capacity design effects as defined by 5.3. In the opposite


case (q < 1,5) the verification shall be carried out using the design seismic combination as defined by 5.5. The verifications may be carried out according to the relevant rules of EN 1993-2 or EN 1994-2 for steel or composite deck respectively. 5.8 Foundations 5.8.1 General (1)P Bridge foundation systems shall be designed to comply with the general requirements set forth in 5.1 of EN 1998-5. Bridge foundations shall not be intentionally used as sources of hysteretic energy dissipation and therefore shall, as far as practicable, be designed to remain undamaged under the design seismic action. (2)P Soil structure interaction shall be assessed where necessary on the basis of the relevant provisions of Section 6 of EN 1998-5. 5.8.2 Design action effects (1)P For the purpose of resistance verification, the design action effects on the foundations shall be evaluated as follows: (2)P Bridges of limited ductile behaviour (q ≤ 1,5) The design action effects shall be those of relation (5.6) with seismic effects obtained from the linear analysis of the structure, under the design seismic combination defined in 5.5, multiplied by the q-factor used. (3)P Bridges of ductile behaviour (q > 1,5) The design action effects shall be obtained by applying the capacity design procedure to the piers according to 5.3. (4) For bridges designed by non-linear analysis or bridges with seismic isolation, provisions of 4.2.4.4 (1)e or 7.6.3 respectively are applicable. 5.8.3 Resistance verification (1)P The resistance verification of the foundations elements shall be carried out in accordance with 5.4.1 (Direct foundations) and 5.4.2 (Piles and piers) of EN 1998-5.


6. DETAILING

6.1 General (1)P The rules of this section are applicable to structures of ductile behaviour and aim at securing a minimum level of curvature/rotation ductility at the plastic hinges. (2)P For structures of limited ductile behaviour rules for the detailing of critical sections and specific non-ductile components are given in 6.5. (3)P In general no formation of plastic hinges is allowed in the deck. Therefore there is no need for the application of special detailing rules other than those valid for bridges under non-seismic permanent and variable actions. 6.2 Concrete piers 6.2.1 Confinement (1)P Ductile behaviour of the compression concrete zone shall be secured within the potential plastic hinge regions. (2)P In potential hinge regions where the normalized axial force (see 5.3.(3)) exceeds the following limit: ηk = ΝEd/Acfck > 0,08 (6.1) confinement of the compression zone according to 6.2.1.3 is in general necessary. (3)P No confinement is required in piers with flanged sections (box- or I-Section) if, under ultimate seismic load conditions, a curvature ductility µΦ = 13 is attainable with the maximum compressive strain in the concrete not exceeding the value of εcu = 0,35% (6.2) (4)P In cases of deep compression zones, the confinement may be limited to that depth in which the compressive strain exceeds 0,5εcu (5)P The quantity of confining reinforcement is defined by the mechanical reinforcement ratio: ωwd = ρw.fyd/fcd (6.3) where: ρw is the transverse reinforcement ratio defined in 6.2.1.1 or 6.2.1.2.


6.2.1.1 Rectangular sections

(1)P The transverse reinforcement ratio is defined as: ρw = Asw/sLb (6.4) where: Asw is the total area of hoops or ties in the one direction of confinement. sL is the spacing of hoops or ties in the longitudinal direction, subject to the following

restrictions: sL ≤ 6 longitudinal bar diameters sL ≤ 1/5 of the smallest dimension of the concrete core. b is the dimension of the concrete core perpendicular to the direction of the confinement

under consideration, measured to the outside of the perimeter hoop. (2)P The transverse distance sT between hoop legs or supplementary cross ties shall not exceed 1/3 of the smallest dimension bmin of the concrete core, nor 200mm. (3)P Bars inclined at an angle α > 0 to the direction of confinement shall be assumed to contribute to the total area Asw of equation (6.5) by their area multiplied by cosα.

6.2.1.2 Circular sections

(1)P The volumetric ratio ρw of the spiral reinforcement relative to the concrete core is used ρw = 4Asp/Dsp.sL (6.5) where: Asp is the area of the spiral or hoop bar Dsp is the diameter of the spiral or hoop bar sL is the spacing of these bars, subject to the following restrictions sL ≤ 6 longitudinal bar diameters sL ≤ 1/5 of the diameter of the concrete core.


4sT2

4sT1

4 closed overlapping hoops

3 closed overlapping hoops + crossties

Typical details using overlapping hoops and crossties

Dsp

≤ 0,6Dsp

Typical detail using interlocking spirals/hoops

≤ min (b min / 3 , 200mm)

Closed overlaping hoops + crossties

bmin

Figure 6.1 Typical details of confining reinforcement

4sT1

sT1 sT2

3sT2

9sT1

4sT2


6.2.1.3 Required confining reinforcement

(1)P The minimum amount of confining reinforcement shall be determined as follows: • for rectangular hoops and crossties

≥ minw,reqw,rwd, 3

2,max ωωω (6.6)

where

ωw,req = cc

c

AA

ληk + 0,13cd

yd

ff

(ρL-0,01) (6.7)

where: Ac is the gross concrete area of the section, Acc is the confined (core) concrete area of the section, and λ factor given in Table 6.1 ρL is the reinforcement ratio of the longitudinal reinforcement Depending on the intended seismic behaviour of the bridge the minimum values given in Table 6.1 are applicable

Table 6.1: Minimum values of λ and ωw,min

Seismic Behaviour λ ωw,min

Ductile 0,37 0,18

Limited ductile 0,28 0,12

• for circular hoops or spirals ( )minw,reqw,wd.c ,1,4max ωωω ≥ (6.8) (2)P When rectangular hoops and crossties are used, the minimum reinforcement condition shall be satisfied in both transverse directions. (3)P Interlocking spirals/hoops are quite efficient for confining approximately rectangular sections. The distance between the centres of interlocking spirals/hoops shall not exceed 0,6Dsp, where Dsp is the diameter of the spiral/hoop (see figure 6.1).

6.2.1.4 Extent of confinement - Length of potential plastic hinges

(1)P When ηk = ΝEd/Ac fck ≤ 0,3 the design length Lh of potential plastic hinges shall be estimated as the largest of the following:


• depth of pier section perpendicular to the axis of the hinge • distance from the point of max. moment to the point where the moment is reduced by

20%. (2)P When 0,6 ≥ ηk > 0,3 the design length of the potential plastic hinges as determined above shall be increased by 50%. (3) The design length of plastic hinges (Lh) defined above is intended to be used exclusively for detailing the reinforcement of the hinge. It must not be used for estimating the rotation of the hinge. (4)P When confining reinforcement is required the amount specified in 6.2.1.3 shall be provided over the entire length of the plastic hinge. Outside the length of the hinge the transverse reinforcement shall be gradually reduced to the amount required by other criteria. The amount of transverse reinforcement provided over an additional length Lh adjacent to the theoretical end of the plastic hinge shall not be less than 50% of the amount of the confining reinforcement. 6.2.2 Buckling of longitudinal compression reinforcement (1)P Buckling of longitudinal reinforcement must be avoided along potential hinge areas even after several cycles into the plastic region. Therefore all main longitudinal bars shall be restrained against outward buckling by transverse reinforcement (hoops or crossties) perpendicular to the longitudinal bars at a maximum (longitudinal) spacing sL = δφL, where φL is the diameter of the longitudinal bars. Coefficient δ depends on the ratio ft/fy of the tensile strength ft to the yield strength fy of the transverse reinforcement, in terms of characteristic values, according to the following relation: 5 ≤ δ = 2,5 (ft / fy) + 2,25 ≤ 6 (6.9) (2) Along straight section boundaries, restraining of longitudinal bars shall be effected in either of the following ways: a. Through a perimeter tie engaged by intermediate corss-ties at alternate locations of

longitudinal bars, at transverse (horizontal) spacing st not exceeding 200 mm. The cross ties shall have 135o-hooks at one end and 90o-hook at the other. The hooks shall be alternated in both horizontal and vertical directions. In sections of large dimensions the perimeter tie may be spliced using appropriate overlap length combined with hooks. When ηk > 0,30, 90o-hooks are not allowed for the cross-ties. In this case it is allowed to use lapped crossed ties with 135o-hooks.

b. Through overlapping closed ties arranged so that every corner bar and at least every

alternate internal longitudinal bar in engaged by a tie leg. The transverse (horizontal) spacing sT of the tie legs should not exceed 200 mm.

(3)P The minimum amount of transverse ties shall be determined as follows: At /sT = Σ Asfys /1,6fyt (mm2/m) (6.10)


where: At is the area of one tie leg, in mm2, sT is the transverse distance between tie legs, in m, ΣAs is the sum of the areas of the longitudinal bars restrained by the tie in mm2, fyt is the yield strength of the tie, and fys is the yield strength of the longitudinal reinforcement. 6.2.3 Other rules (1)P Due to the potential loss of concrete cover in the plastic hinge region, the anchorage of the confining reinforcement shall be effected through 135o hooks surrounding a longitudinal bar plus adequate extension (min. 10 diameters) into the core concrete. (2)P Similar anchoring or full strength weld is required for the lapping of spirals or hoops within potential plastic hinge regions. In this case laps of successive spirals or hoops, when located along the perimeter of the member, should be displaced according to 8.7.2 of EN1992-1. (3)P No splicing by lapping or welding of longitudinal reinforcement is allowed within the plastic hinge region. For mechanical couplers see 5.6.3(2) of EN1998-1. 6.2.4 Hollow piers (1) The following rules are not compulsory in regions of low seismicity. Note: For regions of low seismicity see Note in 2.3.7 (2) Unless appropriate justification is provided, the ratio b/t of the clear width b to the thickness t of the walls, in the plastic region (length Lh of 6.2.1.4) of hollow piers having a single or multiple box cross section, should not exceed 8. (3) For hollow cylindrical piers the above limitation is valid for the ratio Di /t, where Di is the inside diameter. (4) In simple or multiple box section piers and when the ratio ηκ ≤ 0,20, there is no need for verification of confining reinforcement in accordance with 6.2.1, if the provisions of 6.2.2 are met.

6.3 Steel piers (1)P For ductile structures the detailing rules of 6.5, 6.6, 6.7 and 6.8 of EN 1998-1, as modified by 5.7 of the present Part, shall be applied.


6.4 Foundations 6.4.1 Spread foundation (1)P All spread foundations such as footings, rafts, box-type caissons, piers etc., are not allowed to enter the plastic range under the design seismic action, and hence they do not require special detailing reinforcement. 6.4.2 Pile foundations (1)P In the case of pile foundations, it is sometimes difficult, to avoid localized hinging in the piles. Pile integrity and ductile behaviour must be secured in such cases. (2) The potential hinge locations are: (a) at the pile heads, adjacent to the pile cap, when the rotation of pile cap about horizontal

axis, transverse to the seismic action, is very small, due to large stiffness of pile group in this degree of freedom.

(b) at the depth where maximum bending moments develop in the pile. This depth should be

estimated by rational analysis, accounting for the effective pile flexural stiffness (see 2.3.6.1), the lateral soil stiffness and the rotational stiffness of the pile group at the pile cap.

(c) at the interfaces of soil layers having markedly different shear deformability, due to

kinematic pile-soil interaction (see 5.4.2 (1)P of EN 1998-5). (3)P For the location (a) confining reinforcement of the amount specified in 6.2.1.3 along a vertical length equal to 3 pile diameters, shall be provided. (4)P For the locations (b) and (c), unless a more rigorous analysis is performed, longitudinal as well as confining reinforcement of the same amount as that required at the pile head shall be provided for a length of two pile diameters on each side of the point of maximum moment or interface, for cases (b) and (c) respectively. 6.5 Structures of limited ductile behaviour 6.5.1 Verification of ductility of critical sections (1)P For structures of limited ductile behaviour designed with q ≤ 1,5 and located in areas of moderate to high seismicity, the following rules are applicable to the critical sections, aiming at securing a minimum of limited ductility. Note 1: For regions of low seismicity see Note in 2.3.7

Note 2: The National Annex may define simplified verification rules for bridges in low seismicity regions. The application of the same rules as given below is recommended.


(2)P A section is considered to be critical, i.e. location of a potential plastic hinge when: MRd / MEd < 1,30 (6.11) where: MEd is the maximum design moment under the seismic action combinations and MRd is the minimum flexural resistance of the section under the same combination. (3) As far as possible the location of potential plastic hinges should be accessible for inspection. (4)P Where according to 6.2.1 (3), confinement is necessary for attaining a minimum curvature ductility µΦ=7, confining reinforcement as required by 6.2.1.3 for limited ductility, shall be provided. In such a case it is also required to secure the longitudinal reinforcement against buckling, according to the rules of 6.2.2. 6.5.2 Avoidance of brittle failure of specific non-ductile components (1)P Non-ductile structural components such as fixed bearings, sockets and anchorages for cables and stays and other non-ductile connections shall be designed using seismic action effects multiplied by the q-factor used in the analysis or capacity design effects determined from the strength of the relevant ductile elements (cables) and an overstrength factor of at least 1,3. (2)P This verification may be omitted if it can be proven that the integrity of the structure is not affected by the failure of such connections. This proof must also cover the risk of sequential failure, such as may occur in stays of cable-stayed bridges. 6.6 Bearings and seismic links 6.6.1 General requirements (1)P Non-seismic horizontal actions on the deck shall be transmitted to the supporting elements (abutments or piers) through the structural connections which may be monolithic, or through bearings. For non-seismic actions the bearings shall be checked according to the relevant codes and standards (Parts 2 of relevant ECs and EN 1337). (2)P The design seismic action shall in general be transmitted through the bearings. However, seismic links (as defined in 6.6.3) may be used to transmit the entire design seismic action provided that dynamic shock effects are mitigated and properly taken into account. These seismic links should generally allow the non-seismic displacements of the bridge without transmitting significant loads. When seismic links are used the connection between the deck and the substructure shall be properly modeled. As a minimum, a linear approximation of the force-displacement relationship of the linked structure shall be used. Note: Certain types of seismic links may not be applicable to bridges subject to high horizontal non-

seismic actions or to bridges with special displacement limitations, as for instance in railway bridges.


Figure 6.2

Force-displacement relationship for linked structure (3)P The structural integrity of the bridge must be secured under extreme seismic displacements. This requirement shall be implemented at fixed supports through capacity design of the normal bearings (see 6.6.2.1) or through provision of additional links as a second line of defense (see 6.6.3(3)). At moveable connections either adequate overlap (seat) lengths according to 6.6.4 shall be provided or seismic links shall be used. (4)P All types of bearings and seismic links shall be accessible for inspection and maintenance and shall be replaceable without major difficulty. 6.6.2 Bearings

6.6.2.1 Fixed bearings

(1)P The design seismic actions on fixed bearings shall be determined as capacity design effects; (2) Fixed bearings may be designed solely for the effects of the design seismic combinations, provided that they can be replaced without difficulties and that seismic links are provided as a second line of defence.

6.6.2.2 Moveable bearings

(1)P Moveable bearings shall accommodate without damage the total design seismic displacement determined according to 2.3.6.3 (2).

d d y s +s

2

1

Approximation F y

F

s : Slack of the link

dy : Yield deflection of supporting element

1 : Stiffness of bearing

2 : Stiffness of supporting element

2

1

2

1

2


6.6.2.3 Elastomeric bearings

(1) Elastomeric bearings may be used in the following arrangements: a. On individual supports, to accommodate imposed deformations and resist only non-

seismic horizontal actions while the resistance to the design seismic action is provided by structural connections (monolithic or through fixed bearings) of the deck to other supporting elements (piers or abutments).

b. On all or on individual supports, with the same function as in (a) above, combined with seismic links which are designed to resist the seismic action.

c. On all supports, to resist both the non-seismic and the seismic actions. (2) Elastomeric bearings used in arrangements (a) and (b) shall be designed to resist the maximum shear deformation corresponding to the design seismic action. (3) The seismic behaviour of bridges, in which the seismic action is resisted entirely by elastomeric bearings on all supports (arrangement (c) above), is governed by the large flexibility of the bearings. Such bridges and the bearings shall be designed in accordance with Section 7. 6.6.3 Seismic links, holding-down devices, shock transmission units

6.6.3.1 Seismic links

(1) Seismic links may consist of shear key arrangements, buffers, and/or linkage bolts or cables. Friction connections are not considered as positive linkage. (2) Seismic links are required in the following cases: a. In combination with elastomeric bearings, when the links are designed to carry the

design seismic action. b. In combination with fixed bearings which are not designed with capacity design effects. c. Between the deck and abutment or pier, at moveable end-supports, in the longitudinal

direction, when the requirements for minimum overlap length according to 6.6.4, are not satisfied.

d. Between adjacent sections of the deck at intermediate separation joints (located within the span).

(3)P The design actions for the seismic links of the previous paragraph shall be determined as follows: • In cases (a), (b) and (c) as capacity design effects (the horizontal resistance of the

bearings shall be assumed zero). • In case(d), in the absence of a rational analysis taking into account the dynamic

interaction of the deck(s) the linkage elements may be designed for an action equal to 1,5αgQd where αg = ag/g, with αg the design ground acceleration, and Qd is the weight of


the section of the deck linked to a pier or abutment, or the least of the weights of the two deck sections on either side of the intermediate separation joint.

(4)P The links shall be provided with adequate slack or margins so as to remain inactive: • under the design seismic action in cases (c)and (d) • under non-seismic actions in case (a). (5) When using seismic links, means for reducing shock effects should be provided.

6.6.3.2 Holding-down devices

(1)P Holding down devices shall be provided at all supports where the total vertical design seismic reaction opposes and exceeds a minimum percentage of the permanent load compressive reactions.

Note: Minimum percentages may be defined in the National Annex. Recommended values are the following: • 80% in structures of ductile behavior where the vertical design seismic reaction is determined as a

capacity design effect where the plastic hinges have developed their overstrength capacities. • 50% in structures of limited ductile behaviour where the vertical design seismic reaction is

determined from the analysis under the design seismic action only (including the contribution of the vertical seismic component).

(2) The above requirement refers to the total vertical reaction of the deck on a support and is not applicable to individual bearings of the same support. However, no lift-off of individual bearings shall take place under the design seismic combination as defined by 5.5.

6.6.3.3 Shock transmission units (STU)

(1) Shock transmission units (STU) are devices which provide velocity-dependent

restraint of the relative displacement between deck and supporting element (pier or abutment):

• For low velocity movements (v < v1), such as those due to temperature effects or creep

and shrinkage of the deck, the movement is practically free (with very low reaction). • For high velocity movements (v > v2), such as those due to seismic or braking actions, the

movement is blocked and the device acts practically as rigid connection. • Shock transmission units may be provided with a force limiting function, that limits the

force transmitted through the unit (for v > v2) to a defined upper bound Fmax. Note: The order of magnitude of the velocities mentioned above is v1 ≅ 0,1 mm/s, v2 ≅ 1,0 mm/s (2)P The manufacturer of the units shall provide full description of the laws defining the behaviour of the units (force-displacement and force-velocity relations) as well as any influence of environmental factors (mainly temperature, aging, cumulative travel) on this


behaviour. All values of parameters necessary for the definition of the behaviour of the units (including the values of v1, v2, Fmax, for the cases mentioned in (1)) as well as the geometrical data and design resistance FRd of the units and its connections, shall be provided. The above information shall be based on appropriate official test results. (3)P When STUs with force limiting function are used to resist seismic forces they shall have a design resistance FRd not less than: • The reaction corresponding to the capacity design effects, in the case of ductile bridges • The design seismic reaction multiplied by the q-factor used, in the case of limited ductile

bridges. The devices shall provide sufficient displacement capability for all slow velocity actions and full force capacity at their displaced status. (4)P When STUs without force limiting function are used in bridges subject to seismic design situations, the devices shall provide sufficient displacement capability to accommodate the total design value of the relative displacement dEd, as defined by 2.3.6.3 (2) or 7.6.2 (4)P for bridges with seismic isolation. (5)P All STUs shall be accessible for inspection and maintenance/replacement. 6.6.4 Minimum overlap lengths (1)P At supports where relative displacement between supported and supporting elements is intended under seismic conditions, a minimum overlap length shall be provided. (2)P This overlap length shall be such as to ensure that the function of the support is maintained under extreme seismic displacements. (3) At an end support on an abutment and in the absence of a more accurate estimation the minimum overlap length lov may be estimated as follows:

lov = lm + deg + des (6.12) deg = Leffvg/ca ≤ 2dg (6.13) where: lm is the minimum support length securing the safe transmission of the vertical reaction,

but no less than 40cm, deg is the effective displacement of the two parts due to differential seismic ground

displacement, dg is the design value of the peak ground displacement as defined by 3.2.2.4 of EN 1998-1, vg is the peak ground velocity estimated from the design ground acceleration ag, using the

relation vg = 0,16STcag (6.14)


where: ag, S and Tc are as defined in 3.2.2.2 of EN 1998-1. When the bridge site is at a distance less

than 5km from a known seismically active fault, capable to produce a seismic event of magnitude M ≥ 6.5, the values of vg and dg estimated above, should be doubled.

ca is the apparent phase velocity of the seismic waves in the soil. In the absence of more

accurate data the conservative values given in Table 6.2 may be used.

Table 6.2: Estimates of apparent phase velocity ca

Soil Class ca (m/sec) A

B and C D and E

3000 2000 1500

Leff is the effective length of deck, taken as the distance from the deck joint in question to

the nearest full connection of the deck to the substructure. If the deck is fully connected to more than one pier, then Leff shall be taken as the distance between the support and the center of the group of piers. In this contexct “full connection” means a connection of the deck or deck section to a substructure element, either monolithically or through fixed bearing, seismic links, or STU.

des is the effective seismic displacement of the support due to the deformation of the

structure, estimated as follows:

• For decks connected to piers either monolithically or through fixed bearings, acting as full seismic links, des = dEd, where dEd is the total longitudinal design seismic displacement determined according to equation (2.7), 2.3.6.3.

• For decks connected to piers or to an abutment through seismic links with slack equal to s:

des = dEd + s (6.15)

(4) In the case of an intermediate separation joint between two sections of the deck loν, should be estimated by taking the square root of the sum of the squares of the values calculated for each of the two sections of the deck as per (3) above. In the case of an end support of a deck section on an intermediate pier, loν should be estimated as per (3) above and increased by the maximum seismic displacement of the top of the pier dE. 6.7 Concrete abutments and retaining walls 6.7.1 General requirements (1)P All critical structural components of the abutments shall be designed to remain essentially elastic under the design seismic action. The design of the foundation shall be in accordance with 5.8. Depending on the structural function of the horizontal connection between abutment and deck the following rules are applicable. Note: Regarding controlled damage in abutment back-walls see 2.3.6.3 (6).


6.7.2 Abutments flexibly connected to the deck (1) In this type of abutment the deck is supported through sliding or elastomeric bearings. The elastomeric bearings (or the seismic links if provided) may be designed to contribute to the seismic resistance of the deck but not to that of the abutments. (2)P The following actions, assumed to act in phase, shall be considered for the seismic design of these abutments: a. Earth pressures including seismic effects determined according to Section 7 of EN1998-

5. b. Inertia forces acting on the mass of the abutment and on the mass of earthfill lying over

its foundation. In general these effects may be determined using the design ground acceleration ag.

c. Actions from the bearings determined as capacity design effects according to 5.3 (7) and 5.3 (8) if a ductile behaviour has been assumed for the bridge. If the bridge is designed for q = 1,0, then the reaction on the bearings resulting from the seismic analysis shall be used.

(3)P When the earth pressures assumed in (a) above are determined according to EN 1998-1, on the basis of an acceptable displacement of the abutment, adequate provision for this displacement shall be made in determining the gap between the deck and the abutment backwall. In this case it must also be ensured that the displacement assumed in determining actions (a) can actually take place before a potential failure of the abutment itself occurs. For this reason the design of the body of the abutment shall be effected considering the seismic part of actions (a) increased by 30%. 6.7.3 Abutments rigidly connected to the deck (1) The connection of the abutment to the deck is either monolithic or through fixed bearings or through links designed to carry the seismic action. Such abutments have a major contribution in the seismic resistance, both in the longitudinal and in the transverse direction. (2)P The model to be used for the analysis of the bridge must incorporate, in an appropriate way, the soil interaction at the abutments using either realistic values or upper and lower bounds for the relevant stiffness parameters. (3) When the seismic resistance of the bridge is secured by the contribution of other supporting elements as well (piers), the use of high and low estimates for the soil parameters is recommended in order to arrive at results which are on the safe side both for the abutments and for the piers. (4)P A behaviour factor q = 1,5 shall be used, in the analysis of the bridge. (5)P The following actions shall be taken into account in the longitudinal direction:


a. Inertia forces acting on the mass of the structure which may be estimated using the

Fundamental Mode Method (4.2.2.) b. Static earth pressures acting on both abutments (EO) c. The additional seismic earth pressures ∆Εd= Ed – Eo (6.16) where: Ed is the total earth pressure acting on the abutment under seismic conditions as per EN1998-5. The pressures ∆Ed are assumed to act in the same direction on both abutments. (6)P The deck to abutment connection (including fixed bearings or links if provided) shall be designed, in this case, for the action effects resulting from the above paragraphs. Reactions on the passive side may be taken into account according to (8) below. (7) In order that damage of the soil or the embankment behind the abutments is kept within acceptable limits the design seismic displacement should not exceed 6 cm for bridges of γΙ=1,0 where γΙ is the importance factor according to 2.1. (8) The soil reaction activated by the movement towards the fill, of the abutment and of wingwalls monolithically connected to the abutment, is assumed to act on the following surfaces: • In the logitudinal direction, on the external face of the backwall of that abutment which

moves against the soil or fill. • In the transverse direction, on the internal face of those wingwalls which move against

the fill. These reactions may be estimated on the basis of horizontal soil moduli corresponding to the specific geotechnical conditions. The abutment elements concerned should be designed to resist this soil reaction, in addition to the static earth pressures. (9) When an abutment is buried in strong soils for more than 80% of its height, it can be considered as fully locked-in. In that case q = 1 shall be used and the inertia forces determined on the basis of the design ground acceleration (that is without spectral amplification). 6.7.4 Culverts with large overburden (1) When a culvert has a large depth of fill over the top slab (exceeding 1/2 of its span), the assumptions of inertial seismic response used in 6.7.3, lead to unrealistic results. In such a case realistic results may be obtained by neglecting the inertial response and assessing the response caused by the kinematic compatibility between the culvert structure and the free-field seismic deformation of the surrounding soil, corresponding to the design seismic action. (2) The free-field seismic soil deformation may be assumed as a uniform shear-strain field (see Figure 6.3) with shear strain:


s

gs v

vγ = (6.17)

where vg is the peak ground velocity estimated according to 6.6.4, and vs is the shear wave velocity in the soil under the shear strain corresponding to the ground

acceleration. This value may be estimated, from the value vs,max for small strains, using Table 4.1 of EN 1998-5.

6.7.5 Retaining walls (1)P Free standing retaining walls shall be designed according to the rules of 6.7.2 (2) and (3), without any action from bearings.

Figure 6.3 Free-field soil deformation γs

γs/2 γs/2


7. BRIDGES WITH SEISMIC ISOLATION

7.1 Scope (1)P This section covers the design of bridges that are provided with a special isolating system, aiming to reduce their response due to the horizontal seismic action. The isolating units are arranged over the isolation interface, usually located under the deck and over the top of the piers/abutments. (2) The reduction of the response may be achieved: • by lengthening of the fundamental period of the structure (effect of period shift in the

response spectrum), which reduces forces but increases displacements • by increasing the damping, which reduces displacements and may reduce forces • by a combination of the two effects. 7.2 Definitions and symbols (1)P Definitions Isolating system: is the collection of components used for providing seismic isolation, located on the isolation interface. Isolator units or isolators: are the individual components, constituting the isolation system. Each unit provides a single or a combination of the following functions: • vertical–load carrying capability combined with increased lateral flexibility and high

vertical rigidity • energy dissipation (hysteretic, viscous, frictional) • recentering capability • horizontal restraint (sufficient elastic rigidity) under non-seismic service horizontal loads Substructure(s): is the part(s) of the structure located under the isolation interface, usually consisting of the piers and abutments. The horizontal flexibility of the substructures should in general be accounted for. Superstructure: is the part of the structure located above the isolation interface. In bridges this part is usually the deck.

Effective stiffness center: is the stiffness center at the top of the isolation interface, considering the superstructure rigid, but accounting for the flexibilities of the isolator units and of the substructure(s).


Design displacement of the isolating system in a principal direction, is the maximum horizontal displacement (relative to the ground) in the same direction of the stiffness center, occurring under the design seismic action.

Total design displacement of isolator unit i including the effects of mass eccentricities and moment about the vertical axis (see 7.5.4 (5)).

Effective stiffness of the isolating system in a principal direction, is the ratio of the value of the total horizontal force transferred through the isolation interface, concurrent to the design displacement in the same direction, divided by the absolute value of that design displacement (secant stiffness). Effective Period is the fundamental period in the direction considered, of a single degree of freedom system having the mass of the superstructure and stiffness equal to the effective stiffness of the isolating system, as given in 7.6.4. Effective damping of the isolating system is the value of viscous damping, ratio corresponding to the energy dissipated by the isolation system during cyclic response at the design displacement. Normal elastomeric bearings are elastomeric bearings satisfying the requirements of 7.5.2.2b(2). Such bearings may be used without performing the special tests of Annex J. Special elastomeric bearings are elastomeric bearings successfully tested according to the requirements of Annex J. (2)P Symbols See 1.6.6. 7.3 Basic requirements and Compliance criteria (1)P The basic requirements set forth in 2.2 of the present part, shall be satisfied. (2)P The seismic response of the superstructure and substructures shall remain essentially elastic. (3) The structure is deemed to satisfy above basic requirements, if the design is carried out in accordance with the procedures given in clauses 7.4 and 7.5, and meets the requirements set in clauses 7.6 and 7.7.


(4)P Increased reliability is required for the strength and integrity of the isolating system, due to the critical role of its displacement capability for the safety of the structure. This reliability is deemed to be achieved if the isolating system is designed according to the requirements of 7.6.2. (5) All isolator units, with the exception of normal elastomeric bearings to 7.5.2.2b(2), shall be subjected to Qualification and Prototype tests according to Annex J. 7.4 Seismic action 7.4.1 Design spectra (1)P The spectrum used should not be lower than the elastic response spectrum defined in 3.2.2.2 of EN1998-1 applicable for non-isolated structures. (2) Particular attention should be given to the fact that the safety of structures with seismic isolation depends mainly on the displacement demands for the isolating system, that are directly proportional to the value of period TD. Therefore a sufficiently conservative (high) value of TD should be selected.

(3)P Site-specific spectra considering near source effects must be used, when the site is located within a distance of 10 km from a known seismic source that may produce event of Moment Magnitude higher than 6.5 or from a subduction source. 7.4.2 Time-history representation (1)P The provisions of 3.2.2.4 are applicable. 7.5 Analysis procedures and modeling 7.5.1 General (1) The following analysis procedures, with conditions for application specified in 7.5.3, are foreseen by the present part: (a) Fundamental mode spectrum analysis (b) Multimode spectrum analysis (c) Time-history non-linear analysis (2) In addition to the conditions given in 7.5.3, basic prerequisites for the application of methods (a) and (b) are the following:


• The usually non-linear force–displacement relation of the isolating system shall be approximated with sufficient accuracy by the effective stiffness (Keff) i.e. the secant value of the stiffness at the design displacement (see Fig. 7.1). This representation shall be based on successive approximations of the design displacement (ddc).

• The energy dissipation of the isolating system shall be expressed in terms of equivalent

viscous damping as the “effective damping” (ξeff). (3) A special case is presented by an isolating system consisting exclusively of low damping elastomeric bearings (equivalent viscous damping ratio approximately 0,05). In this case the normal linear dynamic analysis methods given in 4.2 are applicable. The elastomeric bearings may be considered as linear elastic elements, deforming in shear (and possibly in compression). Their damping may be assumed equal to the global viscous damping of the structure (see also 7.5.2.2b(2)). The entire structure must remain essentially elastic.

7.5.2 Design properties of the isolating system

7.5.2.1 Stiffness in vertical direction

(1)P The vertical load carrying isolator units shall be sufficiently stiff in the vertical direction. (2) This requirement is considered to be satisfied when the horizontal displacement at the centre of mass of the superstructure, due to the vertical flexibility of the isolator units, is less than 5% of the design displacement ddc. This condition need not be checked if sliding or normal laminated elastomeric bearings are used as vertical load carrying elements at the design interface.

7.5.2.2 Design properties in horizontal directions

(1) The design properties of the isolators depend on their behaviour, which may be one or a combination of the following: a. Hysteretic behaviour (1) The force-displacement relation of the isolator unit may be approximated by a bi-linear relation, as shown in Fig. 7.1, for an isolator unit i (index i is omitted)


(2) (2) The parameters of the bi-linear approximation are the following: dy = Yield displacement Fy = Yield force at monotonic loading F0 = Force at zero displacement at cyclic loading Ke = Elastic stiffness at monotonic loading = Unloading stiffness at cyclic loading Kp = Post elastic (tangent) stiffness ED = Dissipated energy per cycle at the design displacement ddc, equal to the area enclosed by

the actual hysterisis loop b. Elastomeric bearings (1) Elastomeric bearings considered in this part are laminated rubber bearings consisting of rubber layers reinforced by integrally bonded steel plates. With regard to damping, elastomeric bearings are distinguished in low damping and high damping bearings. (2) Low damping elastomeric bearings have an equivalent viscous damping ratio ξ approximately 0,05. Such bearings have a cyclic behaviour similar to that of the hysteretic behaviour with very slender hysteresis loops. Their behaviour may be approximated by that of a linear elastic element. (3) High damping elastomeric bearings show a substantial hysteresis loops, corresponding to an equivalent viscous damping ratio ξ usually between 0,10 and 0,20. Their behaviour should be considered as linear hysteretic. (4) From the point of view of required special tests for seismic performance, elastomeric bearings are distinguished in this part as normal and special.

Figure 7.1: Bilinear approximation of hysteretic behaviour

Ke Keff

Kp F0 Fy

Fmax

dd ED dy


(5) Normal elastomeric bearings are low damping bearings, satisfying all the following conditions: - Equivalent viscous damping: ξ = 0,05 (±20%) - Shape factor: 10 ≤ S ≤ 15 - Unscragged secant shear modulus at shear strain of 2,0: G = 1,0 MPa (±15%) Note: Scragging occurs in elastomeric bearings that are subjected to one or more cycles of high shear

deformation before testing. Scragged bearings show a significant drop of the shear stiffness in subsequent cycles. It appears however that the original (virgin) shear stiffness of the bearings is practically recovered after a certain time (a few months). This effect is prominent mainly in high damping and in low modulus bearings.

(6) Normal elastomeric bearings need not be subjected to the Prototype and other Tests of Annex J. (7) Special elastomeric bearings are elastomeric bearings specially tested according to the requirements of Annex J. c. Viscous behaviour

(1) The force of viscous devices is proportional to αv , where ν = •

d = (d)dtd is the velocity

of motion. This force is zero at the maximum displacement dmax = dd and therefore does not contribute to the effective stiffness of the isolating system. The force-displacement relation of a viscous device is shown in Fig.7.2 (for sinusoidal motion), depending on the value of the exponent α.

d = dd sin(ωt) with ω = 2π/Teff F = Cvα = Fmax (cos(ωt))α Fmax = C(ddω)α ED = λ(α) Fmax dd

( ) ( )( )αΓ

αΓαλ+

+= +

2

2α2 0,51

2

Γ( ) = is the gamma function Figure 7.2: Viscous behaviour d. Friction behaviour (1) Sliding devices with flat sliding surface limit the force transmitted to the

superstructure to Fmax = µdNSdsign(

•

d ) (7.1)

Fmax

dd ED

F

d

α = 1

α < 1


where: µd is the dynamic friction coefficient NSd is the normal force through the device, and

sign(•

d ) is the sign of the velocity vector •

d . Such devices however can result in substantial permanent offset displacements. Therefore they should be used in combination with devices providing adequate restoring force. (2) Sliding devices with spherical sliding surface of radius Rb (e.g. Friction Pendulum bearings) provide a restoring force at the design displacement dd equal to NSddd/Rb. For such a bearing the force displacement relation is:

signNµdRN

F Sdddb

Sdmax += (

•

d d) (7.2)

(3) In both the above cases the energy dissipated per cycle ED (see Figure 7.3), at the design displacement dd amounts to: ED = 4µdNSddd (7.3) (4) The dynamic friction coefficient µd depends mainly on: • the composition of the sliding surfaces • the use or not of lubrication • the bearing pressure on the sliding surface • the velocity of sliding and should be determined by appropriate tests (Annex J).

Rb

Figure 7.3: Friction behaviour

Force

F0 = Fmax= µdNSd

Force

NSd NSd

Fmax Fmax

dd dd

Fmax F0 = µdNSd

Kp = NSd/Rb

ED ED


Note: It should be noted that for lubricated pure virgin PTFE sliding on polished stainless steel surface, at the range of velocities corresponding to seismic motions, and under the usual range of bearing pressures, the dynamic friction coefficient may be quite low (≤ 0,01). Such bearings practically do not offer energy dissipation (Ed ≅ 0).

7.5.2.3 Variability of properties of the isolator units

(1)P The adequacy of the isolating system and the adequacy and design properties of the isolator units shall be determined by the Prototype and other Tests specified in Annex J. The design properties of normal elastomeric bearings satisfying the conditions of 7.5.2.2b(5), may be assumed in accordance with following paragraphs (5) and (6). (2)P The properties of the isolator units, and hence those of the isolating system, may be affected by aging, temperature, loading history (scragging), contamination, and cumulative travel (wear). In addition to the set of nominal Design Properties (DP) derived from the Prototype Tests of Annex J, two sets of design properties of the isolating system shall be properly established, following special tests or according to the provisions of the same Annex: • Upper bound design properties (UBDP), and • Lower bound design properties (LBDP). (3)P Independent of the method of analysis, two analyses shall be performed in general. One using the upper bound design properties (UBDP) and leading to the maximum forces in the substructure and the deck, and another using the lower bound design (LBDP) properties and leading to the maximum displacements of the isolating system and the deck. (4) Multimode spectrum analysis or Time-history analysis may be performed on the basis of the set of the nominal Design properties (DP), only if the design displacements ddc, resulting from a Fundamental mode analysis, according to cl. 7.5.4, based on UBDP and LBDP, do not differ from that corresponding to the DP by more than ±15%. (5) The nominal design properties of normal elastomeric bearings satisfying the conditions of 7.5.2.2b(2), may be assumed as follows:

Shear modulus G = 1,0 MN/m2 Equivalent viscous damping ξeff = 0,05

(6) The variability of the design properties of normal elastomeric bearings, due to aging and temperature (not below 0 oC) may be limited to the value of G and be assumed as follows:


Upper Bound Design Properties Gmax = 1,5 MN/m2 Lower Bound Design Properties Gmin = 0,9 MN/m2

7.5.3 Conditions for application of analysis methods (1)P The Fundamental mode spectrum analysis may be applied when all of the following conditions are met: 1. The distance of the bridge site to the nearest known seismically active fault exceeds 10

km. 2. When the conditions of 4.2.2.2 are met 3. The ground conditions of the site correspond to one of the subsoil classes A, B, C or E of

3.1.1 of EN1998-1 4. When the effective damping does not exceed 0,30. 5. The isolating system has following characteristics: i. The force-displacement relation of the isolation units in each direction is essentially

independent from following effects: • The rate of loading • The magnitude of concurrent vertical load • The magnitude of concurrent horizontal load in the transverse direction

ii. The effective stiffness of the isolating system at the design displacement is at least equal to 1/3 of the stiffness at 1/5 of the design displacement.

(3)P Multimode Spectrum Analysis may be applied when the conditions 3, 4 and 5 of paragraph (1)P are met. (4) Time-history non-linear analysis may be applied for the design of any isolated bridge. 7.5.4 Fundamental mode spectrum analysis (1) The rigid deck model (see 4.2.2.3) may be used in all cases. (2)P The shear force transferred through the isolating interface in each principal direction shall be estimated, considering the superstructure to behave as a single degree of freedom system and using: • the effective stiffness of the isolation system Keff • the effective damping of the isolation system ξeff • the mass of the superstructure Wd/g • the spectral acceleration (see 3.2.2.2 of EN1998-1) corresponding

to the effective period Teff, with ηeff = η(ξeff) Se(Teff, neff)


Effective stiffness Keff = Σ Keff,i (7.4)

Keff,i is the composite stiffness of the isolator unit and the corresponding substructure (pier) i.

Effective damping

= 2

dceff

iD,eff 2

1dK

ΣEπ

ξ (7.5)

where, ΣED,i is the sum of dissipated energies of all isolators i in a full deformation cycle at the design displacement ddc.

Effective Period eff

deff 2

gKW

πT = (7.6)

(3) This leads to the following results for TC < Teff (see also Figure 7.4)

Table 7.1 Max acceleration Smax and design displacement ddc

Teff gmaxS

ddc

TC ≤ Teff < TD 2,5 ⋅eff

C

TT

Sηeffαg ⋅C

eff

TT

dC

TD ≤ Teff 2,5 gff2eff

DC αSηT

TTe C

C

D dTT

⋅

where,

g

γg

α g,475Igg

aa== (7.7)

and 2Ceff2C TSηa

πgd

g0,625= (7.8)

The value of ηeff may be taken from the expression

eff

eff 0,050,10

ξη

+= (7.9)


• Maximum shear force g

SWV max

dd = (7.10)

where,

S, TC and TD are parameters of the design spectrum depending on the soil, according to 3.2.2.2

of EN1998-1,

ag is the design ground acceleration on rock corresponding to the importance category of the

bridge ,

γI is the importance factor of the bridge, and

ag,475 is the design ground acceleration corresponding to a design seismic event with reference return period 475 years.

Note: For a pier of height Hi, having a displacement stiffness Ks,i (kN/m), supported by a foundation with rotation stiffness Kf,i (kNm), and carrying isolator unit i with effective stiffness Ki,i (kN/m), the composite stiffness Keff,i is derived from:

if,

2i

is,ii,ieff,

111

K

H

KKK++=

(4) In essentially non-linear systems, Keff and ξeff depend on the design displacement ddc (see Fig. 7.1). Successive approximations of ddc shall be performed to limit deviations between assumed and calculated values within 5%. (5) For the assessment of the seismic action effects on the isolating system and the substructures, in the principal transverse direction (y), the influence of plan eccentricity in the

TC

TD TC

TB TD

2.5

1,0

TB

TD/TC

1,0

Teff

Teff

Acceleration spectrum

gαSηS

geff

max

Displacement spectrum

C

dc

dd

Figure 7.4: Acceleration and displacement spectra


longitudinal direction etot,x (between the effective stiffness center and the center of mass of the deck) on the displacements of isolator unit i, shall be evaluated by amplifying then design displacement ddc effects, by factor δyi calculated as follows: ddi = δyiddc (7.11)

ix

xtot,yi 1 x

r re

δ += (7.12)

with

( )

yi

yi2i2

x ΣΚKxΣ

r = (7.13)

where ddi is the total design displacement of isolator unit i etot,x = eacc+ ex is the total eccentricity in the longitudinal direction, including the effect of accidental eccentricity eacc = 0,03L L is the total length of the deck between joints r is the radius of gyration of the deck mass about vertical axis through its center of mass xi is the coordinate of isolator unit i relative to the effective stiffness center Kyi is the effective composite stiffnesses of isolator unit and pier i, in the y direction. (6)P The rule of 4.2.1.4 (2) shall be applied for the combination of components of the seismic action. 7.5.5 Multimode Spectrum Analysis (1)P Modeling of the isolating system shall reflect with sufficient accuracy: • the spatial distribution of the isolator units and the relevant overturning effects • the translation in both horizontal directions and the rotation about vertical axis of the

superstructure. (2)P Modeling of the superstructure shall reflect with sufficient accuracy its deformation in plan. Accidental mass eccentricity need not be considered. (3) Modeling of the substructures should reflect with sufficient accuracy the distribution of their stiffness properties and at least the rotational stiffness of the foundation. When the pier has significant mass and height or is immersed in water, its mass distribution should also be properly modeled.


(4) The effective damping

= 2

dceff

iD,eff 2

1dK

ΣEπ

ξ may be applied only to modes having

periods higher than 0,8 Teff. For all other modes, unless a more rigorous estimation of the relevant damping ratio is made, the damping ratio corresponding to the structure without seismic isolation should be used. (5)P The rule of 4.2.1.4 par (2) shall be applied for the combination of the horizontal components of the seismic action. (6) The resulting displacement of the stiffness center of the isolating system (dd,m) and the resulting total shear force transferred through the isolation interface (Vd,m), in each of the two-horizontal directions, are subject to lower bounds defined by following ratios:

0,85fd,

md,d ≥

dd

=ρ (7.14)

0,85fd,

md,v ≥

VV

=ρ (7.15)

where dd,f, Vd,f are respectively the design displacement and the shear force transferred through the isolation interface, calculated according to the fundamental mode spectrum analysis of 7.5.4 (7)P In case the above conditions are not met, the relevant effects on the isolation system, deck and the substructures should be amplified by multiplying with:

d

0,85ρ

for the seismic displacements, or (7.16)

v

0,85ρ

for the seismic forces and moments (7.17)

(8) The limitations of paragraph (6) and the relevant corrections of paragraph (7), need not be applied in cases where the isolating system does not meet the conditions 5i of 7.5.3. They also need not be applied where the bridge cannot be approximated (even crudely) by a single degree of freedom model. Such cases appear in: • bridges with high piers, the mass of which has a significant influence on the displacement

of the deck • bridges with a substantial eccentricity ex in the longitudinal direction between the center

of mass of the deck and the effective stiffness center (ex > 0,10L)


In such cases it is recommended that the limitations and corrections of (6) and (7) are applied in each direction to displacements and forces derived from the fundamental mode of the actual bridge model in the same direction. 7.5.6 Time history analysis (1)P Modeling of the isolating system, in addition to the requirements of 7.5.5 (1), shall reflect with sufficient accuracy any of the following effects, if any of these effects is relevant, for the properties of the isolating system: • the rate of loading • the magnitude of concurrent vertical load • the magnitude of concurrent horizontal load in transverse direction. (2)P Modeling shall conform to the requirements of 7.5.5 (2) and (3). (3)P Paragraphs. (6), (7) and (8) of 7.5.5, are applicable for the results of Time-History analysis also. 7.5.7 Vertical component of seismic action (1) The effects of the vertical component of the seismic action may be assessed by linear response spectrum analysis, regardless of the method used for the assessment of the response to the horizontal seismic action. For the combination of the action effects 4.2.1.4 is applicable.

7.6 Verifications 7.6.1 Design seismic combination (1)P The design combination (5.6) of section 5 is applicable. (2)P The seismic action for the isolating system shall be taken in accordance with 7.6.2, and that for the superstructure and substructure in accordance with 7.6.3. 7.6.2 Isolating system (1)P The required increased reliability of the isolating system (see 7.3.(4)) shall be implemented by designing it for a seismic event with increased design ground acceleration:

ag,IS = γISag. (7.18) where


γIS = 1,50 (7.19) (2)P The increased design ground acceleration defined in the previous paragraph need not exceed that corresponding to a seismic event of 2400 years return period. (3) In the absence of a more accurate assessment, the design effects of such an event on the isolating system may be estimated as follows: • The Fundamental mode spectrum analysis to 7.5.4 shall be carried out, using the

spectrum defined in par. (1) or (2) above, leading to increased values of maximum shear force Vm and design displacement of the effective stiffness center dmc. In this analysis the values of Keff, Teff and ξeff shall correspond to the increased value dmc.

• The increased seismic design forces and total seismic design displacements at each isolator unit, including overturning effects when relevant, shall be calculated from those corresponding to the seismic design action of 7.4, as derived by any of the methods of 7.5, by multiplying with the ratios Vm/Vd and dmc/ddc respectively.

• The total maximum displacements of each isolator unit shall be obtained by adding to the above increased total design seismic displacements, the offset displacements potentially induced by the permanent actions, by the long-term deformations (concrete shrinkage and creep) of the superstructure and by 50% of the thermal action.

(4)P All components of the isolating system must be capable of functioning at the total maximum displacements. (5)P The design resistance of each load-carrying element of the isolation system shall be higher than the increased maximum design force acting on the element at the total maximum displacement. It shall also be higher than the design force caused by wind loading of the structure in the relevant direction. (6) Isolator units consisting of normal elastomeric bearings shall be verified for the effects of the previous paragraphs in accordance with the rules given in Annex K. (7) No lift-off of isolators carrying vertical force is allowed under the design seismic combination with seismic action as per 7.4. (8) Lift-off of individual isolators of a pier under the increased seismic action defined in 7.6.2, may be allowed if it can be proved that such a lift-off has no local detrimental effects on the integrity of the isolators. No lift-off is allowed for elastomeric bearings. 7.6.3 Substructures and superstructure (1)P The seismic internal forces EEA due to seismic action alone, in the substructures and superstructure shall be derived from the results of an appropriate analysis according to 7.5.


(2) The design seismic forces EE due to seismic action alone, may be derived from the forces EEA of (1) above, after division with the q-factor corresponding to limited ductile/essentially elastic behaviour, i.e. FE = FE,A / q with q ≤ 1,50. (3) All structure elements should be verified to have an essentially elastic behaviour according to the rules of 5.6.2 and 6.5. (4) Design action effects for the foundation shall be in accordance with 5.8.2(2)P. 7.7 Special requirements for the isolating system

7.7.1 Lateral restoring force (1)P The isolating system shall provide at the design displacement a restoring force exceeding that corresponding to 50% of the design displacement by at least 0,025Wd where Wd is the weight of the superstructure mass. 7.7.2 Lateral restraint at the isolation interface (1)P The isolating system shall provide sufficient lateral restraint at the isolation interface to satisfy any relevant requirements of other Eurocodes or standards regarding limitation of displacements/deformations under serviceability criteria. Note: This requirement is usually critical for braking action in railway bridges. (2) When sacrificial bracings (fuse system) are used at a certain support(s) in the final bridge system for implementing serviceability displacement restraints between the deck and substructures, their yield capacity should not exceed 40% of the design seismic force transfered through the isolation interface of the isolated structure, at the same support and direction. If this requirement is not met, the serviceability state requirements (excepting fatigue) of the relevant non-seismic codes (EN 1992-2, EN 1993-2 or EN 1994-2), should be satisfied for the elements of the bridge structure, under the loading for which the restraining bracing is designed, when this loading is increased so that the relevant reaction reaches the yield capacity of the bracing. (3) When shock transmission units with force limiting function (see 6.6.3.3) are used for implementing serviceability displacement restraints, the shock transmission units should be included in the model, the verifications and the testing procedure of the isolating system. 7.7.3 Inspection and Maintenance (1)P All isolator units shall be accessible for inspection and maintenance.


(2)P A periodic inspection and maintenance program for the isolating system and all components crossing the isolation interface shall be elaborated by the responsible designer and applied by the Owner. (4)P Repair, replacement or retrofitting of any isolator unit or any component crossing the isolation interface shall be performed under the direction of the responsible entity, and shall be recorded in detail in a relevant report.


ANNEX A (INFORMATIVE)

CHARACTERISTIC SEISMIC EVENT FOR BRIDGES AND RECOMMENDATIONS FOR THE SELECTION OF DESIGN SEISMIC ACTION DURING THE

CONSTRUCTION PHASE A.1 Characteristic seismic event (1) The characteristic seismic event can be defined by selecting an acceptably low probability (p) of it being exceeded within the design life (td) of the structure. Then the return period of the event (tr) is given by the expression: tr=1/(1-(1-p)1/td) (A.1) (2) The characteristic seismic action (corresponding to γI=1,0) usually reflects a seismic event with a reference return period of approximately 475 years. Such an event has a probability of exceedence ranging between 0,10 and 0,19 for a design life ranging between 50 and 100 years respectively. This level of design action is applicable to the majority of the bridges, which are considered to be of average importance. A.2 Design seismic action for the construction phase (1) Assuming that tc is the duration of the construction phase of a bridge and p is the acceptable probability of exceedence of the design seismic event during this phase, the return period trc is given by equation (A.1), using tc instead of td. For the relatively small values usually associated with tc (tc ≤ 5years), equation (A.1) may be approximated by the following simpler relation:

pt

t crc ≅ (A.2)

It is recommended that the value of p does not exceed 0,05. (2) The value of the design ground acceleration agc corresponding to a return period trc, depends on the seismicity of the region. In many cases the following relation offers an acceptable approximation agc / ag,475 = (trc/tro)k (A.3) where: ag,475 is the standard design ground acceleration corresponding to the reference return period tro = 475 years The value of the exponent k depends on the seismicity of the region. As a rule values in the range of 0,30 – 0,40 may be used. (3) Independently from the design seismic actions, the robustness of all partial bridge structures should be secured during the construction phases.


ANNEX B (INFORMATIVE)

RELATIONSHIP BETWEEN DISPLACEMENT DUCTILITY AND CURVATURE DUCTILITY OF PLASTIC HINGES IN CONCRETE PIERS

(1) Assuming that: • the horizontal displacement at the center of mass of the deck is due only to the deformation

of a fully fixed cantilever pier of length L and that • Lp is the length of the plastic hinge developing at the base of the pier the required curvature ductility µΦ of the hinge corresponding to a structure displacement ductility µd, as defined in 2.3.5.2, is:

) λ( λ

µ

ΦΦ

µ0,5-13

1d

y

uΦ

−+== 1 (B.1)

where: λ = Lp/L (2) In reinforced concrete sections (where the curvature ductility is used as a measure of the ductility of the plastic hinge), the value of the ratio λ is influenced by such effects as the reinforcement tensile strain penetration in the adjoining element, the inclined cracking due to shear-flexure interaction etc. The value of Lp given in E.3.2 (5) may be used. (3) When a considerable part of the deck displacement is due to the deformation of other components which remain elastic after the formation of the plastic hinge, the required curvature ductility µΦd is given by the expression µΦd = 1 + f (µΦ - 1) (B.2) where f = dtot/dp, is the ratio of the total deck displacement dtot to the displacement dp, due to the deformation of the pier only and µΦ is calculated from equation (B.1).

Note: If the seismic action is transmitted between deck and pier through flexible elastomeric bearings inducing e.g. a value of f = 5 and assuming that a certain value of e.g. µΦ = 15, would be required in the case of rigid connection between the deck and pier, the required value of µΦd according to equation (B.4) amounts to 71, which is certainly not available. It is therefore evident that the high flexibility of the elastomeric bearings, used in the same force path with the stiff pier, imposes a practically elastic behavior to the whole system.


ANNEX C (INFORMATIVE)

ESTIMATION OF THE EFFECTIVE STIFFNESS OF R. CONCRETE DUCTILE MEMBERS

C.1 General (1) The effective stiffness of ductile concrete components used in linear seismic analysis should be equal to the secant stiffness at the theoretical yield point. In the absence of a more accurate assessment one of the following approximate methods may be used: C.2 Method 1 (1) The effective moment of inertia Jeff of a pier of constant cross section is estimated as follows: Jeff = 0,08 . Jun + Jcr (C.1) where: Jun is the moment of inertia of the cross-section of the uncracked pier Jcr is the moment of inertia of the cracked section at the yield point of the tensile

reinforcement. This is estimated from the expression: Jcr = My/(Ec.Φy) (C.2) in which My and Φy are the yield monent and curvature of the section respectively and Ec is the elastic modulus of concrete. (2) These expressions have been derived from a parametric analysis of a simplified non-linear model of cantilever pier with hollow rectangular and hollow and solid circular cross-sections. C.3 Method 2 (1) The effective stiffness is estimated from the design ultimate moment MRd and the yield curvature Φy of the plastic hinge section as follows: EcJeff = νMRd/Φy (C.3) where: ν = 1,20 is a correction coefficient reflecting the stiffening effect of the uncracked parts of the pier. Curvature at yield Φy may be assessed as follows: Φy = (εsy - εcy)/ds (C.4)


and ds is the depth of the section to the center of the tensile reinforcement εsy is the yield strain of the reinforcement, εcy compressive strain of the concrete at the yield of the tensile reinforcement taking into

account. Strain εcy can be assessed by a section analysis on the basis of εsy and the actual force of the seismic combination NEd. (2) The assumptions of following value for the yield curvature: for rectangular sections of Φy = 2,1 εsy/d (C.5) and for circular sections of Φy = 2,4 εsy/d (C.6) where d is the effective depth of the section give in general satisfactory approximation. (3) The analysis performed on the basis of a value of EcJeff based on an assumed value of MRd needs to be corrected only if the required value MRd,req is significantly higher that the assumed value MRd. When MRd,req < MRd the displacements should be multiplied by the ratio MRd/MRd,req.


ANNEX D (INFORMATIVE)

SPATIAL VARIABILITY AND ROTATIONAL COMPONENTS OF EARTHQUAKE MOTION

D.1 General (1) The characterization of the spatial variability and rotational components of the earthquake action motion shall be carried out by considering the probable contribution of the P, S, Love and Rayleigh waves to the total earthquake vibration and the variability of ground conditions. However, simplified models can be used. These models must obey the condition that the response spectra of the motion at every point shall not be smaller than the corresponding site dependent response spectra multiplied by 0,75. D.2 Variability of earthquake motion D.2.1 Introduction (1) In general, a spatial variability model of the earthquake motion must be used only if there exist certain geological discontinuities or marked topographical features capable to introduce significant variations in the characteristics of the ground motion, or if the length of the bridge is greater than [600m]. (2) Spatial variability means that the motion at different points is different; the difference may be measured by the correlation function. If the correlation function is near unity, the motions are very similar; if the correlation function is near zero, the motions are very different and are considered independent. Spatial variability is very dependent on the frequency band considered; for distances of about 1000 m there is small variability in the low frequency band (f < 1 Hz) and a large variability in the high frequency band ( f > 5 Hz). In general, the spatial variability model shall be such that two points separated by a given distance shall have independent motions in a certain frequency band. If two motions are considered independent, their contributions to the response may be combined by the "square root of the sum of the squares" rule. (3) Under usual conditions the influence of the spatial variability of the earthquake motion upon the maximum values of the structural response is small. Thus, the spatial variability may be simply disregarded or represented by a very idealized model. A possible model for the spatial variability is presented in this annex; however, other models may be used if they respect the condition stated in D.1. D.2.2 Wave propagation (1) The propagation with velocity c of an earthquake vibration ui(t) (i=1,2,3) between a station a (taken as reference) and a station b is given by


d/c)(tur(t)u += aii

bi (D.1)

where b

iu (t) and aiu (t) are the time histories of vibration at points b and a, ri is the ratio of the

wave amplitudes at b and at a and d is the distance between a and b measured along the ray of the wave. The ratio ri is a measure of the dissipation of the vibration (by geometrical spreading or by frictional attenuation) along the distance (d → ∞, r → 0). (2) Let aa

pqS (ù) be the spectral functions matrix at station a; then the spectral density functions matrix at point b and the joint spectral density functions matrices between point a and b are given by ( ) ( )ωω aa

pqqpbbpq SrrS = (D.2)

( ) ( ) ( )ωωω aapqp

abpq Sc/diexprS = (D.3)

( ) ( ) ( )ωωω aapqp

bapq Sc/diexprS −= (D.4)

In the case of Rayleigh waves, particles at the surface of the ground describe elliptical trajectories. Let x1 and x2 be the horizontal and vertical axis and aa

11S (ω) be the power spectral density of the horizontal accelerations. Then the matrix of the spectral density functions for the vector of horizontal and vertical acceleration [ ϋ1 , ϋ2 ]T is given by

( ) ( ) ( )( ) ( ) ][ aa

22aa21

aa12

aa11aa

SSSS

Sωωωω

ω = (D.5)

where ( ) )(ωSρS aa

112bb

22 =ω (D.6) ( ) ( ) )(ωSiexpωS aa

11aa12 ωρ= (D.7)

( ) ( ) )(ωSiexpρS aa

11aa21 ωω −= (D.8)

with ρ representing the ratio of the vertical to the horizontal component (about 1,5 for an elastic half space). (3) To set up a wave propagation model it is necessary to decompose the earthquake vibration into adequate wave trains of P, S, Love and Rayleigh waves with appropriate attenuation and dispersive characteristics; This decomposition depends at least on epicentral distance and focal depth and information on this subject is still not sufficient. Then the matrix of the spectral density functions for all points of the base is obtained by using equations (D.1) to (D.4) for every pair of points of the base.


D.2.3 Simplified model D.2.3.1 Fundamentals (1) The simplified model for the spatial variability (without wave propagation) should be based on an ensemble of independent motions which have nonzero values in only a limited zone of the earth surface. Those independent motions must respect the following rules: • The spatial variability is the same for all components of the motion; • Each independent motion is band limited. The highest frequency in each band shall not

exceed 3 times the lowest frequency; • For each frequency band, a square mesh is defined. The size of the sides of the square is

taken equal to the wavelength corresponding to the lowest frequency of the band. This wavelength should be computed from the average S-wave velocity for the zone within one or two times the length of the bridge. When the size of the sides of the square is greater than five times the length of the bridge, the motion is idealized as a rigid base motion;

• Every node is allocated an independent motion, with the characteristics corresponding to the soil profile at the node;

• The motions on the sides and inside the squares are obtained by a weighted average of the independent motions at the four nodes of the square. The total motion u is given by

u(y1,y2,t) = ∑=

4eii

1i

(t)uα (D.9)

where: αi is the i-th weighting factor;

eiu is the i-th independent motion.

The weighting factors are given by

)πy

)cos(πy

cos(α11

21i = (D.10)

where l is the length of the side, and y1 and y2 are the coordinates of the point under consideration, for a coordinate system whose axes have the same directions as the mesh and whose origin coincides with the i-th node. (2) When the sites of the nodes have the same soil profile, the characteristics of the earthquake motion are the same everywhere and correspond to the soil profile characteristics. D.2.3.2 Response spectra (1) The quantification of the independent motions by response spectra is carried out through

partial response spectra obtained from the site dependent response spectrum and the period interval of the independent motion. Let T1i and T2i be the lowest and highest period of the i-th frequency band. Then the partial response spectra are defined by the following rules (expressed in terms of a trilogarithmic representation):


• The partial response spectra for the period interval T1i, T2i coincide with the site dependent

response spectrum; • The partial response spectra for periods smaller than T1i /2 coincide with the line

representing the peak value of ground acceleration aig for that interval, which can be determined by

)(I)(I TTa 1ia2iaig

−= (D.11)

• The partial response spectra for periods greater than 2T2i coincide with the line representing

the peak value of ground displacement dig for that band, which can be determined by )(I)(Id TT 1id2idig

−= (D.12)

• In the period interval T1i /2,T1i the partial response spectra are defined by a straight line

passing through the value of the partial spectra at T1i /2 and the geometric mean computed from the value aig of peak ground acceleration for the interval T1i ,T2i and the ordinate value of the site dependent response spectrum at T1i ;

• In the period interval T2i ,2T2i the partial response spectrum is defined by a straight line passing through the value of the partial spectra at 2T2i and the geometric mean computed from the value dig of peak ground displacement for the interval T1i, T2i and the ordinate value of the site dependent response spectrum at T2i .

(2) The functions Ia (T) and Id (T) correspond to the indefinite integrals of the acceleration and displacement power spectrum corresponding to the site dependent response spectrum. Those power spectra may be computed by the following approximate relations: Sa = 0,2 ξ Α2T1,4 for T < TB (D.13)

Sa = 6 ξ V2 T-0,74 for TB < T < TC (D.14)

Sa = 300 ξ D2 T-3.1 for TC < T (D.15) where Sa is the acceleration power spectrum, ξ is the value of the damping ratio and A, V and D are the values of spectral acceleration, velocity and displacement. TB and TC are the response spectrum parameters defined by Table 4.1 of EC8: Part 1.1. Note that the displacement power spectrum is given by Sd (T) = (T/2π)4 Sa (T). (3) Where the differential displacements of the bridge foundations can induce substantial stresses in the structure, the values of the earthquake action effects shall be determined by

∑+∑∑=j jj

2

nmmnnm)dk(llρE (D.16)


)ξ(ξrrrξξr

ξξ)rξ(ξrρ 2

n2m

22nm

22nmnm

3/2

mn 4)(8

++++−

+=

4)(11 (D.17)

where E is the value of the earthquake action effect, lm is the effect due to the m-th mode of vibration, kj is the effect due to the j-th independent motion, dj is the assymptotic value of the spectrum for the j-th motion for large periods expressed in displacements, r = ωm / ωn and ξ'm is the value of the viscous damping ratio for the m-th mode of vibration. D.2.3.3 Power spectra (1) The quantification by power spectra of the independent motions in the frequency band f1i, f2i should be carried out by considering partial power spectra with zero values for frequencies lower than f1i and higher than f2i, and coinciding with the site dependent power spectrum in the frequency band f1i, f2i. D.2.3.4 Time history representation (1) The quantification of the independent motions by artificially generated accelerograms should be carried out in accordance with the consistency criteria stated in Clause 3.2.2.4, interpreted as to applied between the ensemble of accelerograms representing the independent motion and the partial response spectrum referred to in the response spectra section. D.2.4 Relative static displacements model D.2.4.1 General (1) In normal bridges and in the absence of a more rigorous assessment , the effects of spatial seismic motion variability may be approximated by applying transient relative static displacements between a reference support point (r) and all other support points (i) of the bridge on the soil. D.2.4.2 Relative static displacements (1) The relative displacements dri, in each direction, between supports (r) and (i) may be estimated (on a basis similar to that given in 6.6.4 (3) for deg) as follows: gagriri 2d/cvXd ≤= (D.18) where: Xri is the horizontal distance of support (i) from the reference support (r) measured in the longitudinal direction of the bridge, dg and vg are respectively the design values of the peak ground displacement and peak ground velocity as defined by 6.6.4 (3), and


ca is the apparent phase velocity of the seismic wave in the soil. In the absence of more accurate data, the values given in 6.6.4 (3) may be used. (2) When the soil class under the supports r and i is not the same, the most unfavorable but mutually consistent values of vg and dg shall be used. For ca, either a weighted average of the relevant values may be estimated, using as weighting factors the proportions of the length of each layer to the total distance Xri, or the lowest ca value may be used. D.2.4.3 Design action effects (1) The design seismic action effects may be estimated as follows:

EEE 2

rx

2

dod += (D.19)

where: Edo are the design action effects assessed ignoring the spatial variability of the seismic motion,

and Erx are the design action effects caused by the displacement vector {dri} according to (D.18),

in which all displacements are in the longitudinal direction. (2) The approximation defined in D.2.4.2 assumes that the relative displacements of the supports in the transverse and vertical directions do not induce significant action effects. This assumption is valid as a rule, because the relevant displacement vectors are -to a large degree- compatible with rigid body motion of the deck. Otherwise a similar approximation may be used for these displacements as well. D.3 Rotational components D.3.1 Introduction (1) The rotational vibrations originate from the spatial derivatives of the transitional components; in consequence, whenever a spatially variable model is used, their inclusion is necessary for the coherency of the model. Moreover, the introduction of the rotational components does not increase significantly the amount of computations needed to perform the analysis. (2) The consideration of the rotational components about the horizontal axis may be important for structures that are both high and stiff; the consideration of the rotational component about the vertical axis may be important for introducing torsion effects in symmetrical structures. In both cases, however, the contribution of the rotational components to the total response is in most cases small, i.e. do not increase the response more than 10%. Thus, even if the rotational components are not quantified very accurately, the resulting errors in the total response are acceptable.


(3) Due to the generally small contribution of the rotational components to the total value of the response, it is generally preferable to use a very conservative but simple model, as adopted in this subclause, than a very sophisticated model, whose parameters cannot usually be quantified without assumptions which are not, to a large extent, susceptible to rigorous justification. It should be pointed out, however, that the adopted model is a purely kinematical model, where it is assumed that displacements are orthogonal to the propagation direction; furthermore the values of the propagation velocity must also be assumed. Thus, it is also aplicable to Love waves. D.3.2 Wave propagation (1) In a direct and orthogonal coordinate system x1, x2 and x3, with axis x3 vertical, the rotations Θi due to a field of displacements uj are given by

∂∂

−∂

∂=

j

k

k

ji 2

1xu

xu

Θ (D.20)

where (i, j, k) is an even permutation of (1,2,3). Consider a wave represented by a displacement field uj(xk) travelling along xk with velocity c without changes in its profile, which is represented by the equation uj(t) = fj(xk-ct) (D.21) where fj is a shape function. The time and space derivatives of uj are:

ct)(xct)(xf

ct

u−∂

−∂−=

∂

∂

k

kjj (D.22)

ct)(xct)(xf

xu

−∂

−∂=

∂

∂

k

kj

k

j (D.23)

From those two expressions it follows that:

t

ucx

u∂

∂−=

∂

∂ j

k

j 1 (D.24)

which shows how to transform time derivatives into space derivatives which may be used with equation (D.20) to obtain the final result:

t

uc

Θ∂

∂−= j

i 21 (D.25)


t

uc

Θ∂

∂−= i

j 21 (D.26)

Θk = 0 (D.27) (2) The quantification of the rotational spectra presented in D.3.3 and D.3.4 assumes that the total earthquake motion is due to S-waves although S-waves cannot exist on the boundaries of elastic solids. In some cases the value to be considered for c is not the propagation velocity of the S-waves in the foundation soil; it is, however, always conservative to take a small value for c, because the amplitude of rotations are proportional to the inverse of c. It is generally conservative to attribute to c the value of the S-wave propagation on the top soil layer because this layer is generally the softest layer and the S-wave velocity is, in general, sufficiently near the lower bound of the phase velocity of Rayleigh and Love waves; although the group velocity may present significantly lower values, this only occurs for limited frequency bands, usually on the low frequency end of the spectrum, where the frequency content of the rotational spectra is smaller. D.3.3 Response spectra (1) The response spectra description of the six components of the earthquake action motion should be constituted by six mutually independent response spectra. Three of those spectra are the site dependent response spectra for the two horizontal components (axes x and y) and the vertical component (axis z) referred to in subclause 3.2.2.2. The rotational response spectra are defined by where:

cT

(T)πSS eθ

ex2,0

= (D.28)

cT

(T)πSS eθ

ey2,0

= (D.29)

cT

(T)πSS eθ

ez2,0

= (D.30)

where:

θez

θey

θex Sand S,S are the rotational response spectra about axes x, y and z,

Se is the site dependent response spectrum for the horizontal components, c is the S-wave velocity, and T is the period being considered.

Note: Mutually independent response spectra indicates that the combination rule of "square root of the sum of the squares" may be applicable.


D.3.4 Power spectra (1) The power spectra description of the earthquake action motion should be constituted by six mutually independent power spectra. Three of those spectra are the site dependent power spectra for the two horizontal components (axes x and y) and the vertical component (axis z). The power spectra for the rotational accelerations are given by

( ) ( )ωω ScωS2

2èx 4

0,98 (D.31)

( ) ( )ωω ScωS θ

y 2

2

40,98 (D.32)

( ) ( )ωω ScωS 2

2θz 2

(D.33)

where: θz

θy

θx Sand S,S are the rotational power spectra about axes x, y and z;

S is the site dependent power spectrum for the horizontal components; c is the S wave velocity; and ω is the frequency being considered. (2) The rotational power spectra defined in this subclause are consistent with the rotational response spectra defined in the previous subclause. D.3.5 Time history representation (1) The time history representation of the earthquake action motion should be consistent, in terms of the criteria stated in Clause 3.2.2.4 with the response spectra representation defined in Clause 3.2.2.2.


ANNEX E (NORMATIVE)

PROBABLE MATERIAL PROPERTIES AND PLASTIC HINGE DEFORMATION CAPACITIES FOR NON-LINEAR ANALYSES

E.1 Scope (1) This Annex provides guidance for the selection of the probable material properties and for the estimation of the deformation capacities of the plastic hinges. Both are intended for use exclusively for the non-linear analyses defined in 4.2.4 and 4.2.5. E.2 Probable material properties E.2.1 Concrete (1)P Mean values fcm, Ecm according to Table 3.1 of EN 1992-1-1 shall be used. (2)P For unconfined concrete the stress-strain relation for non-linear analysis defined in 3.1.5 (1) of EN 1992-1-1, shall be used, with the values of strains εc1 and εcu1 as defined in Table 3.1 of the same code. (3) For confined concrete the following procedure should be used (see figure E.1):

(a) Concrete stress σc:

rccm,

c

xr

xrfσ

+−=

1 (E.1)

where:

fcm,c

fcm

εc1 εc1,c εcu1 εcu,c

Ecm Esec

Confined concrete

Unconfined concrete

Figure E.1 Stress-srain relation for confined concrete

σc


cc1,

c

εε

x = (E.2)

seccm

cm

EEE

r−

= (E.3)

secant modulus cc1,

ccm,sec ε

fE = (E.4)

fcm,c = fcmλc (E.5)

1,2542

cm

e

cm

e −−+=fσ

fσ

λ 7,9412,254c (E.6)

strain at maximum stress

−+= 1

cm

ccm,

ff

ε 510,002cc1, (E.7)

(b) Effective confining stress σe: σe is the effective confining stress acting in both transverse directions 2 and 3 (σe = σe2 = σe3). This stress may be accessed on the basis of the ratio of confining reinforcement ρw, as defined in 6.2.1.1 or 6.2.1.2, and its probable yield stress fym as follows: • For circular hoops or spirals:

ymwe 21 fαρσ = (E.8)

• For rectangular hoops or ties: σe = αρwfym (E.9) α is the confinement effectiveness factor (see 5.4.3.2.2 of EN 1998-1) For bridge pier elements, confined according to the detailing rules of 6.2.1 and having a min dimension bmin ≅ 1,0m, the value α ≅ 1,0 may be assumed.

Note: If, in the case of orthogonal hoops, the values of ρw in the two transverse directions are not exactly equal (ρw2 ≠ ρw3), the effective confining stress may be estimated as e3e2e σσσ = .

(c) Ultimate concrete strain εcu,c


This strain should correspond to the first fracture of confining hoop reinforcement. Unless a more rigorous estimation is made, it may be assumed as follows:

ccm,

suymsccu,

1,4

f

εfρε += 0,004 (E.10)

where: • ρs = ρw for circular spirals or hoops • ρs = 2ρw for orthogonal hoops, and εsu = εum is the mean value of the reinforcement steel elongation at maximum force (see 3.2.2.2 of EN 1992-1-1) E.2.2 Reinforcement steel (1) In the absence of relevant information on the specific steel for the project, following values may be used:

1,15yk

ym =f

f (E.11)

1,20tk

tm =ff

(E.12)

εsu = εuk (E.13) E.2.3 Structural steel (1) In the absence of relevant information on the specific steel for the project, following values may be used:

1,25yn

ym =f

f (E.14)

1,30un

um =ff

(E.15)

where fyn and fun are the nominal values of the yield and ultimate tensile strength respectively.


E.3 Rotation capacity of plastic hinges E.3.1 General (1) In general the rotation capacities of plastic hinge sections, θp,u (see 4.2.4.4(2)c) should be evaluated on the basis of laboratory tests, satisfying the conditions of 2.3.5.2(3), that have been carried out on similar components. The same is valid for the deformation capacities of tensile elements or of plastic shear mechanisms used in excentric structural steel bracings. (2) The similarity mentioned above refers to the following aspects of the components where relevant: • shape of the component • loading rate • relation between action effects (bending moment, axial force, shear) • reinforcement configuration (longitudinal and transverse reinforcement including

confinement), for reinforced concrete components • local and/or shear buckling conditions for steel components (3) In the absence of specific justification based on actual data, the reduction factor γR,p of relation (4.23) may be assumed as γR,p = 1,40. E.3.2 Reinforced concrete (1) In the absence of appropriate laboratory test results, as mentioned in E.3.1, the plastic rotation capacity θp,u, and the total chord rotation θu of plastic hinges (see Figure 2.4) may be estimated on the basis of the ultimate curvature Φu and the plastic hinge length Lp (see Figure E.2), as follows: θu = θy + θp,u (E.16a)

)LL

(LΦ(Φθ 2) ppyuup, −−= 1 (E.16b)

where L is the distance from the plastic joint to the point of zero moment in the pier Φy is the yield curvature Φy

L

Lp Mp

Figure E.2 : Φy and Φu

Φu


For linear variation of the bending moment, the yield rotation θy may be assumed:

3y

yLΦ

θ = (E.17)

(2) Both Φy and Φu should be assessed by means of a moment curvature analysis of the section under the axial load corresponding to the design seismic combination. When εc ≥ εcu1, only the confined concrete core section should be taken into an account. (3) Φy should be evaluated by idealizing the actual M-Φ diagram by a bilinear diagram of equal area beyond the first yield of reinforcement as shows in Figure E.3. (4) The ultimate curvature Φu at the plastic hinge of the element should be taken as:

dεε

Φ csu

−= (E.18)

where d is the effective section depth εs and εc are the reinforcement and concrete strains respectively (compressive strains negative), defined from the condition that either of the two or both have reached the following ultimate values: • εcu1 for the compression strain of unconfined concrete (see Table 3.1 of EN 1992-1-1) • εcu,c for the compression strain of confined concrete (see E.2.1 (3) (c)) • εsu for the tensile strain of reinforcement (see E.2.1 (3) (c)) (5) For a plastic hinge occurring at the top or the bottom junction of a pier with the deck or the foundation body (footing or pile cap), with longitudinal reinforcement of characteristic yield stress fyk (in MPa) and bar diameter ds, the plastic hinge length Lp may be assumed as follows: Lp = 0,115L + 0,0085fykds (E.19)

Φy Φu

Mu

Yield of first bar

Figure E.3 : Definition of Φy


where L is the distance from the plastic hinge section to the section of zero moment, under the seismic action. (6) The above estimation of the plastic rotation capacity is valid for piers with shear ratio

3,0s ≥=dLα (E.20)

For 1,0 ≤ αs < 3,0 the plastic rotation capacity should be multiplied by the reduction factor

3s

sα

)λ(α = (E.21)


ANNEX F (NORMATIVE)

ADDED MASS OF ENTRAINED WATER FOR IMMERSED PIERS (1)P The total effective mass in a horizontal direction of an immersed pier shall be assumed equal to the sum of: • the actual mass of the pier (without allowance for buoyancy) • the mass of water eventually enclosed within the pier (for hollow piers) • the added mass ma of externally entrained water per unit length of immersed pier. (2) For piers of circular cross section of radius R, ma may be estimated as: ma = ρπR2 (F.1) where ρ is the water density. (3) For piers of elliptical section (see Figure F1)with axes 2ax and 2ay and earthquake action at an angle θ to the x-axis of the section, ma may be estimated as: ma = ρπ (ay

2 cos2θ + ax2sin2θ) (F.2)

(4) For piers of rectangular section with dimensions 2ax.2ay and for earthquake action in the x-direction (see Figure F2), ma may be estimated as: ma = kρπay

2 (F.3) where the value of k is taken from the following table (linear interpolation is permitted)

ay/ax k 0,1 0,2 0,5 1,0 2,0 5,0 10,0 ∞

2,23 1,98 1,70 1,51 1,36 1,21 1,14 1,00

2a x

2a y

ax

ay

θ

Figure F1

Figure F2


ANNEX G (NORMATIVE)

CALCULATION OF CAPACITY DESIGN EFFECTS G.1 General procedure (1)P The following procedure shall be applied in general for each sense and for each of the two horizontal directions of the design seismic action: (2)P Step 1: Calculation of the design flexural strengths MRd,h of the sections of the intended plastic hinges, corresponding to the selected sense and direction of the seismic action (AE). The strengths shall be based on the actual dimensions of the cross-sections and the final amount of longitudinal reinforcement. The calculation shall consider the interaction with the axial force and eventually with the bending moment in the other direction, both resulting from the combination G"+"AE where G is the sum of the permanent actions (gravity loads and post-tensioning) and AE is the design seismic action. (3)P Step 2: Calculation of the variation of action effects ∆FC of the plastic mechanism, caused by the increase of the moments of the plastic hinges (∆Mh), from the values due to the permanent actions (MG,h) to the moment overstrength of the sections. ∆Μh = γoMRd,h – MG,h (G.1) where γo is the overstrength factor defined in 5.3. The effects ∆FC may in general be estimated from equilibrium conditions while reasonable approximations regarding the compatibility of deformations are acceptable. (4)P Step 3: The final capacity design effects FC shall be obtained by superimposing the variation ∆FC to the permanent action effects FG FC = FG +∆FC (G.2) G.2 Simplifications (1) When the bending moment due to the permanent actions at the plastic hinge is negligible compared to the moment overstrength of the section (MG,h << γ0MRd,h), Step 2 above may be replaced by a direct estimation of the effects ∆FC from the effects AE of the design earthquake action. This is usually the case in the transverse direction of the piers or in both directions when the piers are hinged to the deck. In such cases the capacity design shear of pier "i" may be estimated as follows:


VC,i = ∆Vi = iE,iE,

Ih,Rd,o VMMγ

(G.3)

and the capacity design effects on the deck and abutments may be estimated from the relation:

E

E,i

C,iC A

ΣVΣV

∆F ≅ (G.4)


ANNEX H (NORMATIVE)

ANALYSIS OF IRREGULAR BRIDGES H.1 General (1) In irregular bridges, the sequential yielding of the ductile elements (piers) may cause substantial deviations between the results of the equivalent linear analysis performed with the assumption of a global force reduction factor q (behaviour factor), and those of the actual non-linear response of the bridge structure. The deviations are due mainly to the following effects: • The plastic hinges appearing first, develop usually the maximum post-elastic strains, which

may lead to concentration of unacceptably high ductility demands in these hinges • Following the formation of the first plastic hinges, in the stiffer elements, the distribution of

stiffnesses and hence of forces may change from that predicted by the equivalent linear analysis. This may lead to a substantial change of the assumed pattern of plastic hinges.

(2) The realistic response of irregular bridges under the design seismic action can, in general, be estimated by means of a dynamic non-linear time-history analysis, appropriately performed according to 4.2.2.4. (3) An approximation to the non-linear response may also be obtained by a combination of an equivalent linear analysis with a non-linear static analysis (pushover analysis). Guidance for the application of this method is given in H.2. H.2 Application of static non-linear analysis (pushover)

H.2.1 Analysis directions, reference point and target displacements (1)P The non-linear static analysis defined in 4.2.5 shall be carried out in the following two horizontal directions: • the longitudinal direction x, as defined by the centres of the two end-sections of the deck. • the transverse direction y, that shall be assumed at right angles to the longitudinal

direction. (2)P The reference point shall be the centre of mass of the deck.

(3)P In each of the two horizontal directions x and y, defined in (1) above, a static non-linear analysis according to 4.2.5 shall be carried out, till the following target displacements of the reference point are reached: • in x-direction (longitudinal) : dT,x = dEx (H.1) • in y-direction: (transverse) : dT,y = dEy (H.2)


where dE,x is the maximum of the displacements in x-direction, at the centre of mass of the deck, resulting from equivalent linear multimode spectrum analysis (according to 4.2.1.3) assuming q = 1,0, for the following combinations of seismic components: Ex “+” 0,3Ey and Ey “+” 0,3Ex. The spectrum analysis shall be carried out using the effective stiffness of ductile elements as defined in 2.3.6.1. dE,y is the maximum of the displacements in y-direction at the same point and under the same conditions as defined for dE,x above. H.2.2 Load distribution (1)P The horizontal load increments ∆Fi,j assumed acting in the direction investigated, at each loading step j on lumped mass Gi/g, shall be taken equal to: ∆Fi,j = ∆fjGiζi (H.3) where: ∆fj is the normalized load portion applied in step j, and ζi is a shape factor defining the load distribution along the structure (2) Unless a better approximation is used, both of the following distributions should be investigated: a) constant along the deck, where for the deck ζi = 1 (H.4)

and for the piers connected to the deck P

ii z

zζ = (H.5)

with zi equal to the height of point i above the foundation of the individual pier and zP equal to the height of the pier P b) proportional to the first mode shape, where

ζi is proportional to the component in the direction investigated of the modal displacement at point i, of the first mode, in the same direction. The mode having the largest participation factor in the direction investigated, shall be considered as first mode in this direction. Especially for the piers, following approximation may be used alternatively

P

iPT,i z

zζζ = (H.6)

where ζT,P is the value of ζ corresponding to the joint connecting the deck and pier P.


H.2.3 Deformation demands (1)P Deformation demands at each plastic hinge shall be verified using condition (4.22) where θEd are the maximum chord rotation demands, when the target displacement is reached (see 4.2.4.4). (2)P In each direction, the total deformation at the first loading step when the two parts of condition (4.22) become equal, at any of the plastic hinges, defines the design ultimate deformation state of the bridge. If, at this state, the displacement of the reference point is less than the target displacement in the relevant direction, the design shall be considered unsatisfactory and shall have to be modified. (3) In the longitudinal direction of an essentially straight bridge the displacements of all pier heads connected to the deck are practically equal to the displacement of the reference point. In this case the deformation demands of the plastic joints can be assessed directly from the target displacement. H.2.4 Deck verification (1)P It shall be verified that no significant yield occurs in the deck before the target displacement is reached (see 4.2.4.4 (d)). (2) In analysis in the transverse direction, yielding of the deck about the vertical axis is considered to be significant when it reaches the reinforcement of the top slab of the deck at a distance from its edge equal to 1/10 of its width or at the junction with a web if it is closer to the edge. (3) In analysis in the transverse direction, the significant reduction of the torsional stiffness of the deck with increasing torsional moments, should be accounted for. In absence of a rigorous assessment, the values given in 2.3.6.1 (4) may be assumed up to 1/3 of the target displacement and may be further reduced by 50% at the target displacement. (4) Lift-off of all bearings of the same support, before the target displacement is reached should be avoided. Lift-off of individual bearings of the same support, before the target displacement is reached, is acceptable, if it has no tetrimental effects on the bearings. H.2.5 Verification against non-ductile failure modes and soil failure (1)P All members shall be verified against non-ductile failure modes, according to 4.2.4.4 (e), using as design actions the force distribution corresponding to the target displacement.


ANNEX J (NORMATIVE)

TESTS AND VARIATION OF DESIGN PROPERTIES OF SEISMIC ISOLATOR UNITS J.1 Scope (1)P The deformation characteristics and damping values of the isolator units used in the design and analysis of seismic-isolated bridges shall be based on the tests described in this Annex. These tests are not intended for use as quality control tests. (2)P The prototype tests specified in J.2 aim at establishing and validating the nominal design properties of the isolator units. These tests in general may be project specific. However, available results of tests performed on specimens of similar type and size and with similar values of design parameters are acceptable. (3) The purpose of the tests of J.3 is to substantiate properties of the isolators, which are usually not project specific. (4) The variation of the design properties of the isolator units due to environmental and time-dependent influences, beyond the nominal values validated by the prototype tests, should be established by special tests. In the absence of such tests, the influences may be estimated according to the provisions of J.4. (5) Quality control tests of isolator units are not within the scope of this Annex. J.2 Prototype tests J.2.1 General (1)P The tests shall be performed on a minimum of two specimens. Specimens shall not be subjected to any lateral or vertical loading prior to prototype testing. (2)P Full size specimens shall be used in general. Reduced scale specimens may be allowed by the responsible authority only when existing testing facilities do not have the capacity required for testing full-size specimens. (3)P When reduced scale specimens are used, they shall be of the same material and type, geometrically similar to the full-size specimens, and shall be manufactured with the same process and quality control. J.2.2 Sequence of tests (1)P The following sequence of tests shall be performed for the prescribed number of cycles, at a vertical load equal to the average dead load, on all isolator units of a common type and size:


T1 Three fully reversed cycles at plus and minus the maximum thermal displacement with a test velocity not less than 0,1mm/min.

T2 Twenty fully reversed cycles of loading at plus and minus the maximum non-seismic

design reaction, at an average test frequency of 0,5 Hz. Following the cyclic testing, the load shall be held on the specimen for 1 minute.

T3 Three fully reversed cycles at the total design seismic displacement. T4 Twenty fully reversed cycles at 1,0 times the design displacement, starting from the

design offset displacement. The twenty cycles may be applied in four groups of five cycles each, with each group separated by idle time to allow for specimen cooling down.

T5 Repetition of test T2 but with the number of cycles reduced to three. T6 If an isolator unit is also a vertical load-carrying element, then it shall also be tested for

one fully reversed cycle at the total design seismic displacement under following vertical loads:

1,2 QG + |∆FEd| 0,8 QG - |∆FEd| where QG is the dead load and ∆FEd is the additional vertical load due to seismic overturning effects, based on peak response

in the design earthquake. (2) Tests T3, T4 and T6 should be performed at a frequency equal to the inverse of the effective period of the isolating system. Exception from this rule is permitted for isolator units that are not dependent on the rate of loading (the rate of loading has as primary effect the viscous or frictional heating of the specimen). The force displacement characteristics of an isolator unit are considered to be independent from the rate of loading, when there is less than 15% difference on either of the values of Fo and Kp defining the hysteresis loop (see Figure 7.1), when tested for three fully reversed cycles at the design displacement and frequencies in the range of 0,2 to 2,0 times the inverse of the effective period of the isolating system. J.2.3 Determination of isolators characteristics J.2.3.1 Force-displacement characteristics (1)P The effective stiffness of an isolator unit shall be calculated for each cycle of loading as follows:

np

np

eff dd

FFK

−

−= (J.1)


where: dp and dn are the maximum positive and maximum negative test displacement, respectively and Fp and Fn are the maximum positive and negative forces, respectively, for units having hysteretic and frictional behaviour, or the positive and negative forces corresponding to dp and dn respectively, for units having viscoelastic behaviour. J.2.3.2 Damping characteristics (1)P The energy dissipated per cycle EDi of an isolator unit i shall be determined as the area of the hysteresis loop of each of the three fully reversed cycles at the total design displacement of test T3 of J.2.2. J.2.3.3 System adequacy (1)P The performance of the test specimens shall be considered as adequate if the following requirements are satisfied: R1 the force-displacement plots of all tests specified in J.2.2 shall have a positive

incremental force-carrying capacity. R2 in test T1 of J.2.2 the maximum measured force shall not exceed the design value by more

than 5%. R3 in tests T2 and T5 of J.2.2 the maximum measured displacement shall not be less than

95% of the design value. R4 in test T3 of J.2.2, the effective stiffness and the energy dissipated in each of the three

cycles shall be between a maximum and minimum nominal design values, defined by the design.

R5 The minimum/maximum effective stiffness ratio measured in the 20 cycles of test T4 of

J.2.2 shall not be less than 0,7. R6 In test T4 of J.2.2 the ratio minED/maxED for the 20 cycles shall not be less than 0,7. R7 All vertical load-carrying units shall remain stable (i.e. with positive incremental

stiffness) during the test T6 of J.2.2. R8 Following the conclusion of the tests, all test specimens shall be inspected for evidence

of significant deterioration, which may constitute cause for rejection, such as (where relevant):

• Lack of rubber to steel bond • Laminate placement fault


• Surface rubber cracks wider or deeper than 70% of rubber cover thickness • Material peeling over more than 5% of the bonded area • Lack of PTFE to metal bond over more than 5% of the bonded area • Scoring of stainless steel plate by marks deeper or wider than 0,5 mm and over

length exceeding 20 mm • Permanent deformation • Leakage

J.3 Other tests J.3.1 Wear and fatigue tests (1)P These tests shall account for the influence of cumulative travel due to displacements caused by thermal and traffic loadings, over a service life to at least 30 years. (2) For bridges of normal length (up to app. 200m) and unless a different value is substantiated by calculation, the minimum cumulative travel may be taken as 2000m. J.3.2 Low temperature test (1)P If the isolator units are intended to be used in low temperature areas, with design low temperature for seismic situation Tmin,sd < 0oC, then a test shall be performed at this temperature, consisting of 3 fully reversed cycles at the design displacement, with the remaining conditions as specified in test T3 of J.2.2. The specimen shall be kept below freezing for at least 2 days before the test. (2) In the test of J.3.1, 10% of the travel shall be performed under temperature Tmin,sd. Note: Values of Tmin,sd should be given in the National Annex. The recommended value is:

Tmin,sd =0,3 To + 0,7 Tmin where, To is the datum temperature and Tmin is the minimum shade temperature defined by EN 1991-2-5.

J.4 Variation of properties of isolating units J.4.1 General (1) The following provisions provide guidance for the estimation of the variation of design properties of isolator units, aiming at the assessment of Upper and Lower Bound Design Properties (UBDP and LBDP) required for the design of the isolating system, according to 7.5.2.3.


(2) In general the UBDP and LBDP should be established by appropriate tests evaluating the influence of following factors on each property: • f1: aging (including corrosion) • f2: temperature • f3: contamination • f4: cumulative travel (wear) In general the design properties of cyclic response influenced by the above factors are (see Fig. 7.1 and Fig. F.3) • The post elastic stiffness Kp • The force at zero displacement Fo (3) The effect of each of the previous factors fi (i = 1 to 4) on each design property, should be evaluated by comparing the maximum and minimum values (maxDPfi and minDPfi) of the design property, resulting from the influence of factor fi, to the maximum and minimum nominal values (maxDPnom and minDPnom) respectively, of the same property DP, as measured by the Prototype tests of J.2. Following ratios should be the established for the influence of each factor fi on the investigated design property.

nom

ffimax, maxDP

maxDPλ i= (J.2)

nom

ffmin, minDP

minDPλ i

i = (J.3)

(4) The effective UBDP used in the design shall be estimated as follows: UBDP = maxDPnom . λU,f1 . λU,f2 … λU,f5 (J.4) with modification factors λU,f1 = 1 + (λmax,fi –1)ψfi (J.5) where, the combination factors ψfi account for the reduced probability of concurrence of the maximum adverse effects of all factors, and should be assumed as given in Table J.1:

Table J.1: Combination factors ψfi Bridge Importance ψfi

Greater than average 0,90 Average 0,70 Less than average 0,60


(5) For the effective LBDP (and relevant modification factors λL,fi) a similar format as that of relations (J.4) and (J.5) above should be used in general, in conjunction with λmin,fi. However for the commonly used elastomeric and friction bearings, it may be assumed in general that:

λmin, fi = 1 (J.6)

and therefore LBDP = minDPnom (J.7) J.4.2 λmax-values for elastomeric bearings (1) In the absence of appropriate test results, the λmax-values given in following tables J.2 to J.5 may be used for estimation of the UBDP.

Table J.2 f1: Aging

λmax, f1 for Component Kp Fo LDRB 1,1 1,1 HDRB1 1,2 1,2 HDRB2 1,3 1,3 Lead core - 1,0

with following designation for the rubber components: LDRB : Low damping rubber bearing with shear modulus larger than 0,5 MPa HDRB1 : High damping rubber bearing with ξeff ≤ 0,15 and shear modulus larger than

0,5 MPa HDRB2 : High damping rubber bearing with ξeff > 0,15 and shear modulus larger than

0,5 MPa Lead core : Lead core for Lead core rubber bearings

Table J.3 f2: Temperature

λmax, f2 for Kp Fo

Design Temperature Tmin,sd (oC) LDRB HDRB1 HDRB2 LDRB HDRB1 HDRB2

20 1,0 1,0 1,0 1,0 1,0 1,0 0 1,3 1,3 1,3 1,1 1,1 1,2

-10 1,4 1,4 1,4 1,1 1,2 1,4 -30 1,5 2,0 2,5 1,3 1,4 2,0

Table J.4

f3: Contamination λmax,f3 = 1,0


Table J.5 f4: Cumulative travel

λmax,f4 = 1,0 J.4.3 λmax-values for sliding isolator units

(1) In the absence of appropriate test results, the λmax-values given in the following tables J.6 to J.9 may be used for the estimation of the maximum force at zero displacement Fo corresponding to the UBDP. The values given for unlubricated PTFE cover also Friction Pendulum bearings.

Table J.6 f1: Aging

λmax,f1 Component Unlubricated

PTFE Lubricated

PTFE Bimetallic Interfaces

Environment Sealed Unsealed Sealed Unsealed Sealed Unsealed Normal 1,1 1,2 1,3 1,4 2,0 2,2 Severe 1,2 1,5 1,4 1,8 2,2 2,5

Table J.7

f2: Temperature Design

Temperature λmax,f2

Tmin,sd (° C) Unlubricated PTFE

Lubricated PTFE

Bimetallic Interfaces

20 1,0 1,0 0 1,1 1,3

-10 1,2 1,5 -30 1,5 3,0

To be established

by test

Table J.8

f3: Contamination λmax,f3

Installation Unlubricated PTFE

Lubricated PTFE


Sealed, with stainless steel surface facing down

1,0 1,0 1,0

Sealed, with stainless steel surface facing up 1,1 1,1 1,1

Unsealed, with stainless steel surface facing down

1,2 3,0 1,1


Table J.9 f4: Cumulative travel

λmax, f4 Cumulative Travel (km)

Unlubricated PTFE

Lubricated PTFE


≤ 1,0 1,0 1,0 To be

established by test

1,0 < and ≤ 2 1,2 1,0 To be

established by test

(2) Above data are valid under the following conditions:

• Table J.6

- Stainless steel sliding plates are assumed - Unsealed conditions are assumed to allow exposure to water and salt - Severe environment include marine and industrial conditions - Values for bimetallic interfaces apply to stainless steel and bronze interface.

• Table J.8 - Sealing of bearings is assumed to offer contamination protection under all serviceability

conditions.


ANNEX K (NORMATIVE)

VERIFICATION OF ELASTOMERIC BEARINGS UNDER SEISMIC DESIGN SITUATIONS

K.1 General (2) The provisions of this Annex are intended for the verification of laminated elastomeric

bearings used in general under design seismic situations of bridges, including the case of elastomeric bearings used as components of a seismic isolation system. These provisions do not cover non-seismic design situations.

(3) Rules are given for the verification of both Normal and Special elastomeric bearings (in

the sense of 7.5.2.2b). K.2 Components of shear strain

K.2.1 Total design shear strain (1)P The total design shear strain (εtd) shall be determined as the sum of the following components: εtd = εc + εs + εα (K.1) where: εc is the shear strain due to compression εs is the shear strain due to the total design seismic displacement according to 2.3.6.3 and εα is the shear strain due to angular rotation K.2.2 Shear strain due to compression (1)P The shear strain due to compression shall be determined as follows:

Gσ

Sε e

c1,5

= (K.2)

where: G is the shear modulus of the elastomer σe is the maximum effective normal stress of the bearing calculated as:

r

Sde A

Nσ = (K.3)

where: NSd is the maximum axial force on the bearings resulting from the design seismic load

combination Ar is the minimum reduced effective area of the bearing calculated as follows:


• for rectangular bearings with steel plate dimensions bx and by (without holes) Ar = (bx - dEdx) (by - dEdy) (K.4) • for circular bearings with steel plate of diameter D

( )4

2

rDsinδδA −= (K.5)

with /D)arccos(dδ

Ed2= (K.6)

and ddd 2Edy

2EdxEd

+= (K.7) In the above equations dEdx and dEdy are the total relative displacements under seismic conditions, in the x- and y- directions respectively, of the two bearing faces, including the design seismic displacements (with torsional effects) as well as the displacements due to the imposed deformations of the deck (i.e. shrinkage and creep where applicable and 50% of the design thermal effects). S is the shape factor of the relevant elastomer layer, defined as the ratio of the effective

compressed area divided by the side area free to bulge i.e.

• for rectangular bearings: iyx

yx

2 )tb(bbb

S+

= (K.8)

• for circular bearings:

i4tDD = (K.9)

In the above equations ti is the thickness of the elastomer layers. K.2.3 Shear strain due to the total seismic design displacement (1)P The shear strain due to the total design seismic displacement ddd

2Edy

2EdxEd += ,

including torsional effects, shall be determined as follows:

t

Eds t

dε = (K.10)

where: tt = Σti is the total thickness of the elastomer.


K.2.4 Shear strain due to angular rotations (1)P The shear strain due to angular rotations shall be determined as follows: • For rectangular bearings: εα = (bx

2αx + by2αy) / 2ti tt (K.11)

where: αx and αy are the angular rotations across the bx and by dimensions of the bearings respectively. • For circular bearings of diameter D: εα = D2α / 2ti tt (K.12)

where ααα 2y

2x +=

(2) Normally in bridges the influence of εα is negligible for the seismic verification. K.3 Design criteria for elastomeric bearings K.3.1 General (1)P Adequate seismic links shall be provided where it is necessary to secure the structural integrity of the bridge in accordance with 6.6.1 (3). The links shall be designed in accordance with 6.6.3. (2)P The bearings shall comply with the seismic design criteria of K.3.2, K.3.3, and K.3.4 K.3.2 Maximum values of shear strains

Maximum allowable values of shear strains εc, εs, and εtd are given in Table K1.

Table K1 Maximum allowable values of shear strain

Shear Strain Symbol Expression Maximum Value

Due to compression only εc (K.2) 2,5 Due to total design seismic displacement εs (K.10) 2,0

Total design shear strain εtd (K.1) 6,0


K.3.3 Stability (1)P Either of the following criteria must be satisfied bmin / tt ≥ 4 (K.13) or

mint

mine

32

St

bGσ

≤ (K.14)

where: bmin is the minimum dimension of the bearing Smin is the minimum shape factor of the bearing layers tt is the total elastomer thickness and σe is the maximum effective normal stress of the bearing according to K.2.2. K.3.4 Fixing of bearings (1) For normal elastomeric bearings, friction may be considered to insure that sliding of the bearing does not occur, if both the following criteria are satisfied under the most adverse seismic design condition.

e

f0

Ed

Ed

σk

kNV

+≤ (K.15)

σe ≥ 3,0 N/mm2 (K.16) where: k0 is 0,10 for bearings with external elastomeric layer, or 0,50 for bearings with steel plates having

external identations kf is 0,6 for concrete and 0,2 for all other surfaces, VEd and NEd are respectively the shear and the axial force transmitted simultaneously through the bearing according to the design seismic combinations, and σe is the effective normal stress in N/mm2 (expression) (2) When one of the conditions of the previous paragraph is not satisfied, or when special elastomeric bearings are used, positive means of fixing shall be provided, in both bearing sides, to resist the entire maximum design shear force VEd. The shear resistance of concrete under dowel action on both the supported and supporting elements should be verified using appropriate failure models.