Fatigue Behaviour of Orthotropic Steel Decks

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Com mission of the European Com munities

t e c h n i c a l s t e e l r e s e a r c h

Proper t ies and serv ice per formance

Measurement and interpretation of dynamicloads in bridges

Phase 3Fatigue behaviour of orthotropic steel decks

Synthesis Report

EUR 13378 EN



3

Commiss ion of the European Communit ies

t e c h n i c a l s t e e l r e s e a r c h

Properties and service performance

Measurement and interpretation of dynamicloads in bridges

Phase 3Fatigue behaviour of orthotopic steel decks

Edited by:

A. Bruls

Service 'Ponts et Charpentes'Université de Liège

Quai Banning 6B-4000 Liège

Contract No 7210-KD/119/201 /317/411 /609/807(1 July 1986 to 3 1 December 1988)

Synthesis report

Directorate-GeneralScience, Research and Development

1991

PARI [UROP Biblioth.

N.C. EUR 13378 E N



Published by theCOMMISSION OF THE EUROPEAN COMMUNITIES

Directorate-GeneralTelecommunications, Information Industries and Innovation

L-2920 Luxembourg

LEGAL NOTICENeither the Com mission of the European Communities nor any person acting

on behalf of the Commission is responsible for the use which might be made ofthe following information

Cataloguing data can be found at the end of this publication

Luxembourg: Office for Official Publications of the European Communities, 1991

ISBN 92-826-0532-9 Catalogue number: CD-NA-13378-EN-C

© ECSC-EEC-EAEC, Brussels • Luxem bourg, 1991

Printed in Belgium



SUMMARY.

This research, carried out with the financial help of the ECCS, concerned thefatigue strength of orthotropic steel decks of road bridges. It followed twophases that were concerned with the collection of traffic data and measurement

of stresses produced in bridges. Fatigue tests under constant and variableamplitude were carried out on stiffener-plate connections, stiffener-stiffenerconnections with U and V shapes, and stiffener cross-beam connections. Fromthe tests results and calculations some conclusions can be drawn which aredirectly usable in bridge design. However, some unexpected behaviour occuredand some connections need more investigation.

Résumé.

Cette recherche, réalisée avec l'aide financière de la CECA, concernait la

résistance à la fatigue des dalles orthotropes de ponts-routes. Elle faisaitsuite à deux phases qui se sont concentrées sur la collecte de donnéesrelatives aux charges du traffic et aux contraintes produites dans les ponts.Les essais de fatigue sous amplitude constante et variable ont été réaliséssur les assemblages raidisseur en U-tôles, les assemblagesraidisseur-raidissèur en U et en V, les assemblages raidisseurs-entretoise.Les résultats des essais et les calculs ont permis de tirer des conclusionsdirectement applicables au calcul des ponts-routes. Néanmoins descomportements imprévisibles s'étant manifestés, certains assemblages demandentdes investigations complémentaires.

Zusammenfassung.

Dieses Forschungsvorhaben wurde mit finanzieller Unterstützung der EGKSdurchgeführt und betraf das ErmUdungsverhalten orthotroper Platten vonStrassenbrüchen. Es folgte auf zwei Phasen, in denen hauptsachlich Daten überVerkehrslasten sowie Beanspruchungen von Brücken gesammelt wurden.Ermüdungsversuche unter konstanten und variablen Amplituden wurden fürfolgende Verbindungen durchgeführt : U-Längsverstelfung/Blech, U-undV-Längsversteifung, und Längsversteifung/Quertrager. Die Ergebnisse derVersuche und der Rechnungen erlaubten Schussfolgerungen, die direkt für dieAuslegung von Brücken anwendbar sind.

Aufgrund unerwarteter Verhaltenswesen der untersuchten Bauteile sind jedochzusatzliche Forschungen notwendig.



This report is a synthesis of the final reports performed

by each Laboratory that has participatedto the common research :

1. H.LEHRKE, Fraunhofer Institut fUr Betriebsfestigkeit, Bartningstrasse H7 _

6100 Darmstadt - Germany. [1]

2. A. BRULS, E. POLEUR, Service "Ponts.et Charpentes", Université de Liège,

6, quai Banning - 1)000 Liège - Belgique. [2]

3. A. BIGNONNET, I.R.S.I.D. 78105 Saint-Germain-en-Laye - France, and

J. CARRACILLI, B. JACOB, L.C.P.C., 58, Bd. Lefebvre - 75732 Paris - France.

C3]

k. S. CARAMELLI, P. CROCE, M. FR0LI, L. SANPAOLESI, Istituto di Scienza delle

Costruzioni, Università di Pisa, Via Diotisalvi, 2 - 56126 Pisa - Italy.

cu

5. H. KOLSTEIN, J. DE BACK, Stevin Laboratory, Universiteit van Delft, 2628 CN

Delft, Nederland. [5]

6. C. BEALES, Transport and Road Research Laboratory, Old Wokingham Road,

Crowthorne, United Kingdom [6].

Coordination : A. BRULS

IV



CONTENTS

Page

SUMMARY 111

RESUME 111

ZUSAMMENFASSUNG 111

1. INTRODU CTIO N 1

1.1 Orthotropic steel deck 1

1.2 Details tested 1

2 . METHODOLOGY OF FATIGU E PRE DICTION 5

2.1 Traffic loads and effects on bridg es 5

2.2 Classical life calculation 6

2.3 The fracture mechanic approach 9

3. CONN ECTION STIFF EN ER -P LATE 12

3.1 Types of connection 123.2 Stress determination 123.3 Fatigu e testing and results 153.4 Fatigu e life calculation 19

3.5 Crack propagation and lifetime calculations 203.6 Conclusion 224. CONNECTION STIFFE NER- STIFFE NER 404.1 Types of connections 404.2 Stress determination 404.3 Test results of the University of Pisa 424.4 Test results of the T.U. DELFT 454.5 Comparison with other research programs 514.6 Conclusions 51

5. CONNE CTION STIFFEN ER- CROS SBE AM 70

.5.1 Types of connection 70

5.2 Stress determination 705.3 Test results of TRR L 725.4 Test results of the T.U. Delft 755.5 Test results of the L.B.F . 77

5.6 Conclusions 79

6. OR THO TRO PI C DECK TO CRO SS BE AM CON NE CTIO N 106

7. APPLICATIONS 111

8. CONCLUS IO NS 112

BIBLIOGRAPHY 115



1. INTRODUCTION.

1.1. ORTHOTROPIC STEEL DECKS (Fig. 1.1).

Orthotropic steel decks are used in bridges with long spans and

in movable bridges in which dead weight has to be as low as possible.

The upper part of these decks is composed of a plate on which the

traffic runs. This plate is covered by a thin surfacing (7 to 12 mm) or an

asphalt surfacing (MO to 70 mm).

Longitudinal stiffeners are welded to the under part of the deck

plate, approximately 300 mm. apart, parallel to the direction of the

traffic lanes. They are usually closed sections with trapezoidal or "V"

shapes although open sections are sometimes used. The stiffeners transmitloads to crossbeams to which they are connected.

Crossbeams are normally spaced at 3 to 5 meters and are connected

to main girders or diaphrams.

Orthotropic steel decks are very sensitive to fatigue damage

because they are directly subjected to the actions of wheel loads which

give rise to stress ranges which are high compared to the dead load

stresses especially if the surfacing is thin.

In order to understand the fatigue behaviour of orthotropic steel

decks, the path of the traffic loads must be considered. Under the action

of wheels, the deck plate acts as a beam on elastic supports (the stiffener

webs). The elasticity of the supports decrases with the spacing of the

crossbeam. The web of the stiffeners are subjected to normal forces

(support reaction) and to bending moment if stiffeners have closed

sections.

Longitudinally the stiffeners are subjected to shear forces,

bending and torsion moments. The butt-welds connecting the longitudinal

stiffeners and the stiffener to crossbeam connection bear these effects.

1.2. Details tested.

This research concerns the study of the fatigue strength of

orthotropic decks with closed stiffeners. Bibliographical researches and

bridge examinations made during the two first phases have enabled us to

define details which are most sensitive to fatigue damage (Fig. 1.1).

- 1 -



* detail 1

* detail 2

* detail 3

*d e t a i l H

stiffener to deckplate connection,

stiffener to stiffener connection,

stiffener to crossbeam connection,

bolted connection crossbeam-orthotropic plate.

For each detail, the stress field was studied by calculation

and/or by measurement in the laboratory. Afterwards constant amplitude

fatigu e tests were carried out to define the S- N cu rv es. F inally specimens

were tested under variable amplitude load cycles simulating traffic

effects.

Distribution of the work between laboratories was as follows :

* detail 1 : stiffener to deck plate connection :

- University of Liège analysed local stresses with the help of

a finite band program and tested the connection under constant

and variable amplitude loading ;

- I R S I D carried ou t constant amplitu de tests on this connection

with a different welding procedure to that in the University of

Liège.

- L.C.P .C. developed and applied a fracture mechanics model.

* D etail 2 : stiffener to stiffener connection :

- Univ ersity of Pisa measured and calculated stresses on a full

size orthotropic deck ;

- University of Liège calculated stress histograms produced by

traffic loads using the results of the 1st and 2nd phases ;

- U niversity of P isa and T.U. Delft tested different designs of

this detail.



* Detail 3 : Stiffener to crossbeam connection.

- T.R.R.L. measured stresses in a full size orthotropic deck and

in an actual bridge ;

- L.B.F. calculated stresses in the web of the crossbeam withthe help of a finite element program.

- University of Liège calculated stress histograms produced by

traffic loads , using the results of the 1st and 2nd phases.

- T.R.R.L, T.U. Delft and L.B.F. tested different designs of this

detail.

* Detail 1J : bolted connection crossbeam-orthotropic plate.

- University of Liège tested this detail.

Using the traffic load data from the 1st and 2nd phases and the

fatigue data from the current research, it has been possible to make

some important conlusions about the fatigue behaviour of orthotropic

steel decks.



Longitudinal deck plate butt weld

Deck plate Transverse deck plate butt weld

/

de ta il ^ Longitudinal stiffener to deck plate Longitudinal stiffenerto crossbeam weld

Longi tud ina ls t i f f e n e r

de ta il (5 ) Longitudinal stiffener splice welds

Crossbeam to deck plate weld de tai l © st if fe ne r

to crossbeam connection(a l te rnat ive connect ions)

Crossbeam Diaphragm

Fi gu re 1 .1 Main welded connect ions in a typ ica l o r th ot op ic br idge deck



2. METHODOLOGY OF FATIGUE PREDICTION.

2.1. Traffic loads and effects on bridges.

During the first and second phase of the research [7][8],

measurements of traffic loads and stresses were carried out on 1 -4 highwaybridges. Different types of bridges and spans from 13 to 1000 m were

examined, most of them included orthotropic steel decks.

The traffic on these bridges was recorded and the number and types of

commercial vehicles identified and classified in a uniform scheme. Axle

loads, were measured by weighbridges, and axle spacings, separation between

vehicles and lateral wheel positions were also recorded. Additional sensors

simultaneously recorded the stresses produced in structural components like

main and cross girders, longitudinal stiffeners and deck plates.

The recorded stress histories were analysed using Level-crossing

and Rain-flow cycle counting methods. The resulting spectra were assessed

in terms of their fatigue damage potential with reference to constant

stress range fatigue design curves from existing or draft codes. Special

attention was paid to the possibility of calculating by means of finite

elements, and to the temperature dependent effect of the asphalt surfacing

in reducing the stresses.

Based on the experimental data a method of computer simulation oftraffic loads and stresses induced was developed and tested. It has proven

a suitable tool not only to reproduce the experimental data, but also to

derive more general results in that any uniform or hypothetical traffic

situation may be studied for the bridges investigated or for any components

of bridges described in terms of influence lines.

The main conclusions were :

1. The level of the measured axle and vehicle loads are much higher than

the allowed loads. Despite this observation, the measured stresses are

never higher than 90 N/mm2 ; which is not very much.

2. The number of vehicles is so high for some traffic, that exclusing the

vibration effect, the number of cycles produced during a life of 100a

years may reach 10° • which is very high.



3. The recording of a traffic flow comprising about 10 0 0 0 axles higher

than 10 kN will be sufficient to establish a spectrum reliably and to

allow a fatigue calculation.

Tw o types of v ehicle, the articulated 2- axled trailer with a 2- axled

semi trailer or with a 3 _axled semi trailer, have been found to produce

about 40 to 50 % of the total fatigue damage on Dutch bridges, although

they represent only 6 to 18 ? of the total number of commercial

vehicles.

i\ . As steel decks support hig h and frequent local stress ranges caused by

wheel loads, a main part of the work was concentrated on orthotropic

steel decks.

During the 2nd phase of the research, the most aggressive traffic was

recorded in the Netherlands on the Rheden bridge, where the frequency of

articulated lorries was hig hest. Therefore, in this work the Rheden traffic

was mainly considered for the determination (by simulation) of the stress

spectra used in fatigue tests. During the period of the third phase, new

measurements of traffic flow w ere made (independently of this w ork ) in

F rance, Germany and I taly. These data su g gest that the load of the vehicles

has not changed very much, but the number of articulated lorries and the

number of loaded lorries has increased. The results obtained by using the

R heden traffic are u sable, but the number of cycles are higher for somerecent traffic [9 ].

It is clear that the development of steel bridges needs the knowlegde of

the fatigue behaviour of the details of an orthotropic deck under variable

amplitude load and a high number of cycles ; thu s it needs the fatigu e

design curve for stress ranges beyond 2.10 cycles.

2. 2. Classical fatigue life calculation.

2 .2.1. Miner's rule.

During the life of a bridge, its components are subjected to

stress spectra induced by traffic. For the fatigue calculation it is

necessary to translate the actual stress spectra by a stress range

histogram.

Cycle counting methods such as Rain-flow or Range-Pair are commonly used to

produce stress range histrograms ; mean stress is not u sually considered.

A bridge is thus influenced by variable stress ranges, occuring randomly,

from the vehicle loads running on the bridge.

The fatigu e behav iour of the details are characterised by S- N

curves.

- 6 -



Fatigue damage is calculated using Miner's rule

D = Z (Ü L)A O l -0 N.

with n. = number of cycles in the stress range histrogram corresponding to

Aa. measured or calculated during a time t.

N, = number of cycles corresponding to Aa. in the S-N curve.

That is the number of cycles with a stress range of Aa.

producing failure.

If D < 1 : no failure

D - 1 : failure

The remaining fatigue life is — g — t.

In Eurocode the fatigue classification of a detail is defined by the value

Aa corresponding to N =2 .10 cycles.C c

Eurocode 3 considers SN curves with two slopes (Fig. 2.1) :

N.Aa3 = este = 5.106. Ao 3 i f Aa £ AaQ

5 6 5N.Aa = este - 5.10 . Aa if Ao^ > Aa i Aa T

D D L

N « » if Aa. > AaLi

where,

Aan corresponds to N =5.10 cycles ;u o g

AaTcorresponds to N = 1 0 cycles.Li L

The fatigue damage corresponds to

A aD n. . Aa5, " n. . Ao \

D E = Z U (_i L_) + E (_j__J_)Aa=Aa L 5.106.Aa 5

A a A o 5. 1 06. A a 3

1 D i D D

The lower part of the curve (m = 5) is important in bridges which

experience large numbers of low stress cycles.

- 7



2.2.2. Existing codes.

To design a structure, designers have to find information in codes

about actions, strengths and calculation methods. For the fatigue

assessment of bridges most of the codes propose the Miner's rule as thecalculation method. The action (traffic) is idealised by one or a set of

standard vehicles and the strengths are presented in the form of S-N

curves.

A. Action.

The action to consider in the fatigue assessment of bridges is the

action of traffic. During the first and second phases of the research,

measurements were made of traffic including the frequencies of vehicle

types and the level of loading of lorries. The traffic were compared to

try to detect local, regional or international influences.

It is not possible to define a spectrum taking all these influences into

account.

In existing codes actions are defined by one, two or three of the

following load models :

- one vehicle moving on an influence line allows the calculation of the

highest stress range (NBN 5 [10], BS 5^00 [11]) ;

- one vehicule moving on an influence line allows the calculation of the

stress range histogram (.La., n. ) by a cycle counting method (BS 5*100).

- a set of vehicles moving on the influence line allows the calculation of

the stress range histrogram (Aai, ni) by a cycle counting method (BS

5100).

B. Strength.

The different SN curves existing in codes are presented inFig. 2.1. The main remarks are :

- for constant amplitude loading, the S-N curves often have a slope of

-1/3 (m = 3) with an endurance limit. Previously the endurance limit

corresponded to N = 2.10 ,now it is associated with N - 5.10 (NBN)

[10], EC3 [12]) or 107(BS 5^00 [11], NEN [13]).

- for variable amplitude loading, the S-N curve is continued beyond the

endurance limit with a slope corresponding to m (NBN), m+2 (BS 5^00) or

2 m -1 (EC3).

o

- EC3 proposes a cut-off value at N - 10 below which the variable

amplitude loading does not produce any fatigue damage.- 8 -



- If the maximum value of the stress range A o m a x is below the endurance

limit, it is assumed that there is no fatigue damage (NBN).

C. Calculation.

A first verification, which is simple and conservative, is to

calculate the maximum stress range for the detail under the load model.

This value is compared with the endurance limit (NBN 5) or with charts

of limiting stress which vary according to the bridge span and detail

class (BS 5^00). The calculated stress range must be below the limiting

stress for the detail to be acceptable by this method.

The second verification is more precise. It involves a fatigue

damage calculation on the basis of the (Ao, n ) histogram defined by ii i the load model, using the Miner's rule.

2.3. The fracture mechanic approach.

The only method suitable for bridge design requires a Miner's

calculation, but the crack growth evaluation in an existing bridge

needs a fracture mechanic approach.

The classical Miner's calculation is very simple and does not take

into account the changes in the structure when a crack is propagating, or the mean stress level onto which the stress variations are

superimposed. Hence this model is very sensitive to the choice of the

S-N curves. For road bridges in which the stress variations are

generally much smaller than the permanent stresses, the computed

lifetimes are highly dependent on the high endurance end of the S-N

curves.

In the fracture mechanic approach, the crack propagation is

computed for each stress cycle, taking into account the present state

of the structure and hence the time history of the stress variations.

There are numerous laws describing this evolution ; the Paris one has

been chosen because of its simplicity and its ability to describe the

crack propagation in this type of structure. The crack propagation

speed is written as :

da m — — ■ c ÅK t where c and m are parameters depending on the

material and AK the stress intensity factor, depending on the stress range and on the local

geometry of the detail.

- 9 -



In order to account for a threshold, under which no crack may be propaged,

the Paris law is modified as

da ra'

_ = c' (AK - AK )dN s

The threshold plays a similar rule than the classical fatigue limit.

If the stress are both positive and negative, only tension is considered

for the crack propagation. But in the real bridges, it is generally assumed

that the residual stresses induce high forces, and the whole stress

variation are considered.

The crack propagation time is obtained by integrating this formula from the

initial crack length a to its .length a at the failure. The number ofo rstress cycles involved is :

1 ar m'N = N + _ ! ƒ (AK - AK ) da

C' a

O

where N is the number of cycles for the crack initiation. The failure

criteria adopted here is : a 0.5e, if e is the plate thickness ; it

corresponds to the loss of rigidity of the structure. The problem is then

to compute the AK values.

The cracked beam theory shows that AK may be generally written as :

AK = /la.Ao.f(a), where f(a) is a function of the structure geometry and

Aa is the stress range in the uncracked section studied.

For the orthotropic deck stiffener-plate connections, the following formula

may be adopted :

AK = /HaCA F + 2a_ A, F + _* 1 A„ F„ + ^ ì L A F + Ü _ A„ F„]o o f 1 1 2 2 2 3 í 3 3 8 ^ ^

where the F. are amplification factors depending on the crack geometry but

independent on the loading case. The A. are the coefficients of the

smoothing polynominal of the stress diagram.

This model has been applied by the LCPC to the stiffener-plate connection

behaviour in the chapter 3.

10



s.ioV

BS5¿00

•-N(log)

Ao. (log)

N(log)

NEN 2063

Ao (log)

AOc AOn .

AOL

,

^"""""^IT

2.1

i

"■ "■ "■»« ̂

0 * S.1

■ " ^ « « ^

2m-1

0 * 1

cut-off

o'

E C 3

AT 4 (log)

cut-off

2.10' S.V* 10*

EC 3

-Nllog)

Figure 2.1 : S-N curves

11 -



3. CONNECTION STIFFENER-PLATE.

3.1. Types of connection.

3.1.1. General considerations.

The main parts of the stiffener-deck plate connection are the deck

plate, the stiffener and the weld between them.

Usually the deck plate is between 10 or 12 mm thick and the

surfacing thickness is around 10 mm or around 60 mm depending of the

material-.

The stiffeners studied are closed trapézoïdal or "V" shaped.

Stiffener dimensions are about 300 mm wide, 250 mm high and 6 mm thick.

They are typically placed 300 mm apart. Because of the closed section of

stiffeners, welding is carried out on only one side of the stiffener web.

Other types of open stiffeners that were used in earlier bridges are not

considered in this research.

3.1.2. Welding prodecures.

Welding procedures have evolved : originally the procedure was

manual metal arc (tests carried out by Maddox [1*1] and JANSS [15]), now it

is often automatic submerged arc welding. Procedures are still evolving toobtain a smaller lack of penetration (tested in this work).

Several tests have been carried out at IRSID to optimize the welding

procedure. The influence of the edge preparation has been checked and

comparisions made between edges chamfered at 60° or 45° and without

chamfer. The welding energy was adjusted to minimize the lack of

penetration.

From several sets of welding tests it has been shown that without edge

preparation and without chamfer, a satisfactory penetration (lack of penetration £ 1 mm) is obtained for a welding energy of 20 Kj/cm, (figure

3.1), even with a gap of 2 mm between the top of the stiffener web and the

deck plate.

3.2. Stress determination.

3.2.1 . Measurements and calculations.

Stresses in the deck plate to stiffener connection are induced by

the direct effects of the wheel load, the orthotropic steel deck behaves

as a beam on elastic supports (stiffener webs) (Fig. 3.2.).

12 -



During the two first phases, the stresses near the weld were

measured under traffic loads and with a test vehicle of known axle load.

The stresses were influenced by the temperature and by the distribution of

wheel loads, both in magnitude and transverse position. The temperature

affects the stiffeners of the surfacing and its composite action with the

steel plate. Another factor is that it is not possible to measure a

stress at the crack initiation point.

In order to have a general approach to the behaviour of the welded

connection under a wheel load, it is necessary to develop a stress

calculation method. Stresses have been calculated by a finite band

program. The frame used for the calculation has the geometry of

orthotropic decks found in Belgian bridges (Fig. 3.3.a.). The points where

stresses are calculated are located in the neighbourhood of the weld (Fig.

3-3.b) ; points A' , 3' : in the deck plate ; points C' , D' : in the weld.

Axial and bending stresses are calculated in the cross section in which

they are the highest (section 0 - axis 1 - Fig. 3.3.a).

With the finite band program it is possible to study the influence of the

following parameters :

- longitudinal location of wheel : longitudinal influence lines are drawn

in Fig. 3.11. It appears that at points A' and B' (deck plate) stress values change sign at a certain distance from section 0. Thus the stress

amplitude at point A' and 3' is higher than the maximum stress obtained

when the wheel is on axis 1 ;

- transverse location of wheel : results are presented in the form of

transverse influence lines (Fig. 3.5 and 3-6).

■ surfacing thickness : two surfacing thicknesses are considered :

- deck without surfacing : load is not distributed ;

- deck with a 60 mm surfacing thickness : load distributes through the

thickness at an angle of 1)5° ; no composite effect is considered.

- dimensions of wheel contact area : different sizes of wheel are studied.

13



3.2.2. Stress histograms.

The calculation of the stress histograms used for the variable

amplitude tests were made using the simulation program of Liège [2][7]

[ 8 ] . For each axle of the traffic the program chose at random a transverse

position on the deck plate corresponding to a value of the transverse

influence line. This v alue, multiplied by the load of the axle gave the

stress induced.

The data introduced in the simulation program are :

1. Traffic : the vehicle axles of the Rheden traffic were divided into

four groups according to their wheel type.

2 . Transverse distribtion : the transverse distribution of the vehicles

was obtained from measurements made during the 1st and the

2nd phases.

3. Transverse influence lines : calculated in section 3.2.1.

The histogram obtained is presented in table 3.1 and figu re 3.7..

It was used for the variable amplitude tests.

3.2.3. Equivalent stress range.

To compare the variable amplitude test results with the constant

amplitude ones, the applied stress spectra were analysed according to

Miner's rule in two different ways, each using an S-N curve with a slop of

- 1 / 3 .

a. An equivalent stress range La was calculated in a way that n-cycles of

that stress range have the same fatigue damaging potential as n-cycles

of the stress spectrum, using a third power relationship ;

*° a "(-T7T z

n . Â o . ' ) 1 / 3 MPae E n. i l

b . The equivalent values proposed by the University of Liège [25] in the

report of the 2 d phase [2 5 ], corresponds to the centre of gravity of

the damage distribution (see Fig. 3.7) ;

Z n < Å o <Ao - _ _ L _ _ MPa

3I niAoi

nm

3I niAoiJ

A°m3

in which : Ao^ - individual stress range

rU - individual number of cycles corresponding to AojAora " equivalent stress range

nm - number of cycles belonging to Aom

This definition is not much influenced by a cutting of the high numberor low cycles which produces not much damage.

In this report both method are used.

14



3.3. Fatigue testing and results.

3.3.1. Test specimens.

The tests presented in this section are :

* previous tests : - W .I . (Maddox) [11]

- CRIF (Janss) [15]

* tests of the University of Liège [2]

* tests of I.R.S.I.D [ 3 ] .

The loading mode and the g eometrical characteristics of the

specimens are giv en in figu re 3 .8. and Table 3.2 .

The material used is of the type E 36—^.

General welding procedure characteristics are :

- no edge stiffener preparation- horizontal position

- one run

- no preheating

- no postheating.

- Specific conditions :

W. I . : manual arc w elding

C.R.I.F . : manuel arc w elding

I.R.S .I.D. : automatic w elding (submerged ar c) .U .Lg. : automatic w elding.

The geometrical characteristics of the welds are given in table 3.2.

3 .3 .2. Test resu lts presentation.

D epending on the stress distribu tion in the deck plate and in the

trough, as well as the weld quality (penetration, throat, thickness,

undercut, ...) fatigue cracking may occur either :

a) from the weld toe on the deck plate, point A fig. 3.9., developing in

the deck plate.

b) from the weld root in the stiffener, point D fig. 3.9., developing in

the throat of the weld.

To determine the stress distribution in the specimen, static tests

were carried out. It is noted that in the median axis of the specimen the

loading is biaxial w ith o2/o1 - I /3 (o1 is in the direction of the bending

stresses). This phenomenon has not been taken in account in the S-N curvesbecause designers do not consider biaxial stresses.

15



From these tests the nominal stresses to be used in the fatigue

S-N diagrams were derived (figure 3.9).

- extrapolation of the stress in the deck plate to the weld-toe, Ao.

definition generally used when cracks propagate through the deck plate

from the weld toe,

- extrapolation of the stress in the trough to the weld root Ao , S

definition generally used when cracks propagate through the weld from

the root of the weld.

Four different failure criteria have been used :

N » crack detection by strain gauge ;

N =■ first visible crack ;

N = measurable charge in stiffness of the specimen, or 25 mm. long crack.

N. = end of test.

The results of the tests made at IRSID and at the University of

Liège are given in tables 3.3. Tests with failure in the deck plate and in

the root weld are considered separately.

3.3.3. Failure in the deck plate Aorf.

3.3.3.1. Constant amplitude tests :

These tests determine the fatigue strength of the deck plate at the weld

toe. Results are given in Table 3.3.a and plotted in Fig. 3.10.

Main conclusions are :

- There are no significant differences between specimens with a 2 mm gap

between the top of the sti f f ener web and the deck plate and specimens

with no gap, provided the lack of penetration is less than 2 mm.

The mean Wöhler curve is determined with R = -1 and R 0 - 0,1 in the

weld root for the specimen tested at IRSID.

-1/3 Aod - 26777 N (m = 3 imposed) .

As the standard deviation is 24 N/mm2, the characteristic value for

97,5 ? is Ao = 163 N/mm2 for N - 2 10 6 cycles, de c

- The two experiments conducted at IRSID show a lower fatigue resis-

tance at R = 0,1 than R » -1. However, two similar tests performed

at Liège do not indicate such a detrimental effect.

16



In only two U.Lg. specimens did failure occur in the deck plate

corresponding to R » 0 (only tensile).

3.3.3.2. Variable amplitude tests :

The loading histogram used for these tests is the stress spectrum calcu

lated in section 3.2.2. (table 3.I- and Fig. 3.7). This histogram simulates

traffic effects. Loads were applied at random to the test specimen.

Two tests were carried out. They are plotted in Fig. 3.10 at their

equivalent values (Ao , n ) calculated from the histogram. Them m

equivalent values correspond to the centre of gravity of the damage

distribution, see figure 3.7.

Results are similar to those obtained with constant amplitude loading.

3.3.I. Failure in the weld Ao .s

3.3.1.1. constant amplitude tests.

Tests results are given in Table 3.3.b and plotted in Fig. 3.11.

Main conclusions are :

- the fatigue strength increases significantly when using automatic

welding, this technique allows larger penetration and throat

thickness at the weld.

- The mean S-N curve is determined :

Ao - 17258 N~ 1 / 3 and

the characteristic value for

97,5 $ is Ao - 111 N/mm2 for N - 2.10 cycles,se c

- The tests show the importance of R ratio. To obtained the failure

in the weld it was necessary to have more tension than compression

at the root of the weld :

- 1 < R < 0 in the ULg specimens

R » 0,1 in the IRSID specimen.

3.3.1.2. variable amplitude fatigue tests :

The loading histogram used for these tests is the stress spectra

calculated in Section 3.2.2. modified to obtain failure in the weldRs - -0,5 instead of - 2,0 at the root for the highest stress range.

- 17 -



The results are plotted in Fig. 3.11. at their equivalent values

corresponding to the centre of gravity of damage distribution.

Results are similar to those obtained under constant amplitude

loading.

3.3.5. Comparison with previous research.

Results from the Welding Institute [1*1], CRIF [15 ], University of

Liège [2 ], and I R SI D [3] are compared in Fig u re 3.11 in terms of nominal

stress range in the trough at the weld root L a (any other representative3

stress could be used) versus the number of cycles to failure of the trough

to deck plate connection M .

The main difference is that in the Liège University, WI and CRIFexperiments cracks initiate at the weld root and the failure occurs in the

weld, while in the IRSID experiments cracks initiate at the weld toe in

the deck plate and the failure occurs in the deck plate.

The results of the tests at IRSID and at the University of Liège

are given in tab. 3-3.

As shown in tables 3.2., the specimens tested by CRIF and by

Maddox, welded by manual arc welding, confirm the importance of lack of

penetration. It is clear from figure 3.11 that the fatigue resistanceincreases significantly when using submerged arc welding, this technique

allowing larger penetration and throat' thick ness of the w eld. Nevertheless

it is shown that cracks initiate at the weld root even with a lack of

penetration of 2 mm. However if the welding operation is properly

optimised, the lack of penetration can be. limited below 1 mm. and, for

alternate or repeated tensile bending in the deck plate, the cracks

initiate at the weld toe, and the fatigue resistance is improved.

f.

V

\

/y

18 -



3.H. Fatigue life calculation.

The characteristic stress range deduced from the fatigu e tests, defined

for N » 2.10 cycles following Eu rocode 3, are :c

Aa — 114 N /mm

2

, if the crack occurs in the w eld (point D) ;scAa. - 163 N/mm2 if the crack occurs in the plate (point A ) ,

dc

If the traffic composition measured at Rheden is considered, the number of

lorries required to cause failure may be calculated (see 3 .2 ). The results

are given in table 3.4.

Putting there data in perspective, the traffic flows recorded on highways

during the 1st and 2 nd phases generally comprise between 100 0 and 40 0 0

lorries during a working day. Such flows produce, after 100 years, between

20 and 80 .10 lorries.The following comments may be made :

1. The fatigue life calculated in the deck (Ao ) is always a little higher

than in the weld (Aa ) : the hig her fatig u e strength is partially

offeset by higher stress ranges produced by the traffic loads ;

2 . A surfacing of 6 0 mm gives a fatig ue life 3 to 5 times longer than the

unsurfaced deck ;

3. An increase in the deck plate thickness of 1 mm gives a fatigue life

around twice as long.

4. An increase in the thickness of the stiff ener redu ces the fatigu e life

a little.

5. The g iven fatigu e lives are pessimistic, because the transverse

position of the traffic flow considered in the calculation is in the

most damaging position.

6. The given fatigu e lines are too pessimistic for surfaced decks where

the calculation doesn't take in to account a composite effect, that

exist mainly by low temperature.

We may conclude that the required thickness of the deckplate depends

of the expected number of lorries, and the thickness of the surfacing.

- 19



3.5. Crack propagation and lifetime calculations.

The fracture mechanics model has been applied to the stiffener-

plate connection of various structures :

- the IRSID test specimens,- the two temporay bridges of Montlhery and Choisy-le-Roi,

- the bridge of Caronte.


The calculation of the crack propagation lifetime for the test

specimens presented in § 3.3. was made by the LCPC to calibrate and check

the model. The tested and computed lifetimes are compared in table 3.5.

Various stress levels have been considered from 150 to 3*40 MPa and the

mean stress described by the ratio R presented in § 2.3. was given either

by R = -1 or R = 0.1. The lifetime calculation was made for 2 initial

crack lengths : 0,05 and 0,H mm.

The very good welding conditions (automatic optimised welding),

are likely to lead to small initial defects and the initial crack length

adopted here is between 0,05 and 0,1 mm. In existing bridges, it is

generally assumed that this length may rise to between 0,3 and 0.5 mm. The

agreement between the tests and the model predictions are fairly good,

given the usual uncertainties in such experimentation.

3.5.2. Montlhery bridge.

In this case the stress diagram is fitted in this case by the

polynomial :

o(z) = 7.533 - 2.706 z + 0.481 z 2 - 0.0788 z s + 0.0055 zk

The lifetimes are calculated using the measured stress histograms,

recorded in 1978 during the first phase of this study. Results are given

in the tables 3-6 and 3.7. In order to compare the results with the

classical Miner's approach, the lower limits of the stress intensity

factor AK are chosen as functions of the fatigue limits (at 5 million

cycles) of the S-N curves : AK = f (a , a ) , following the last formula

of §.2.3.

For this bridge the fairly poor welds may lead to the adoption of

well class iJ5 or 50 MPa and the lifetimes calculated from the fracture

mechanics model will be longer than the calculated using Miner's rule.

For good welds (classes between 63 and 71 MPa), both models give the same

-20



both models give the same results for this stress distribution. The sensi

tivity of the fatigue life to the weld class is much higher in the Miner's

model than in the fracture mechanics one ; the latter seems to be more

realistic. In any case, the short lifetimes found here show that this

bridge was not designed for such a heavy and dense traffic.

3.5.3. Bridge of Caronte.

The stress diagram here is fitted by the polynomial :

a(z) - 6.0211 - 3.287 z + 1.075 z 2 - 0.190 z3

* 0.0123 z**

Table 3-7 shows the computed lifetimes for various R and AK . The stress

variations are those measured in the second phase of this research. Due to

the low stress variations, no damage is expected for classes above 71 MPa

and the lifetimes are always very long. With the Miner's rule, thelifetimes become short for the lowest classes...

3.5.1. Bridge of Choisy-le-Roi.

The stress diagram is represented by the same polynomial as for

Montlhery, because the structure is identical and the loading similar. In

this case two traffic flows have been used for computing the stress time

history by the LCPC's program CASTOR. One is the existing traffic on the

bridge, before an increase forecast after the opening of a new motorwaysection, and the other was assessed in Angers (RN 23), representing the

future traffic on the bridge. The influence surface of the transverse

stress in the deck plate along the stiffener-plate connection is given in

figure 3.12.

Tables 3.8 and 3.9 give the fatigue lives computed for both types

of traffic, various AK .The influence of the initial crack length can beS

seen. The choice of AK = 2 and a - 0.3 mm. seems to be realistic. Under

s 0the existing traffic the fatigue life is approximately 200 years, while

under the predicted traffic it falls to 1)2 years (assuming well class 63

MPa). The results obtained by Miner are very close for this class, but

again the calculations are very sensitive to the fatigue limit. Figure

3.13 shows the crack propagation versus time, for various AK and for thes

two traffic flows. It is clearly seen that the crack length (and the

damage) does not grow as a linear function of time or number of cycles, as

predicted by the Miner's law.

21 -



3.6. Conclusion.

The specimens used in the experiments were welded with automatic

submerged arc welding in an industrial situation. The first step in this

work consisted of optimising the welding parameters. It has been shown thatfull penetration, (lack of penetration less than 1 mm) can nearly be

achieved without edge preparation.

In these conditions, there was no signifiant difference in fatigue

behaviour for specimens with a 2 mm gap between the top of the stiffener

web and the deckplate and those with no gap.

For alternate bending (which best represents the loading in

bridges) with a lack of penetration not greater than 1 mm., the cracks

initiated at the weld toe in the deck plate. Work at the University of

Liège with similar specimnes but with a lack of penetration of 1,5 to 2,5

mm. lead to crack initiation at the root of the weld if the tensile

stresses were higher than the compressive stresses (a m a x > /a /) at this

point. Therefore if the lack of penetration is not greater than 1 mm., it

is possible to exclude cracks in the weld, and to assess the risk of crack

initiation at the weld toe.

The test results allow the required plate thickness to be

determined, depending of the expected lorry traffic flow, and the thickness

of the surfacing (see table 3.1)). It is now possible to choose between the

thickness of the surfacing and the thickness of the deck plate in order to

increase the fatigue life.

22



Stresses

0 .i

Û«/isaa )

70 - SO60 - 7050 - 6040 - 5030 - 4020 - 3010 - 200 - 100

-10 / -0-20 / -10-30 / -20-40 1 -30-50 / -¿0-60 / -50-70 / -60-50 / -70-90 / -SO-100 / -50-110 / -100-120 / -110-130 / -120-140 / -130

TOTAL

;

Number

1319612226351175"27515176

| 12627

6629¿348234316681C685703651S98320

972

40000

Stress-ranges

ha . (N/sra2)

10 - 2020 - 3030 - 4040 •- 5050 - 6060 - 7070 - SOSO - 9090 - 100

100 - 110110 - 120120 - 130130 - 140140 - 150150 - 160

160 - 170170 - ISOISO - 190190 - 200200 - 210210 - 220-

-

Number

6767416324S91650109087260144131425314213481431721117522

19150

TABLE 3.1. : Stress and stress-range hist ograms

at point D' (variable amplitude tests)

- 23



T A B L E 3 . 2 G E O M E T R I C C H A R A C T E R I S T I C S OF T E S T S P E C I M E N S

1

ro

i

GEOMETRIC CHARACTERISTICS(nun)

- deck plate thickness td

- stiffener thickness ts

- stiffener width B

b

- stiffener height h

- location of the load L

- location of the support S,

S2

- type of welding

- gap

- lack of penetration

- fillet weld

W . I .11

6,35

305

150

230

Manuel

5 to 6

4,5 to 5

CRIF

12

6

300

109

250

335

337

Manuel

< 0,5

3,0 to 4,5

3,5 to 4,5

U.LG.

12

6

3Ò0

109

250

85

335

337

Autom.

< 0,5

1,5 to 2,5

4,7 to 6

IRSID

12

6

322

212

226

85

335

335

Autom.

0 or 2

15,5 to 6,5



IO Ol

Lab.

IRSID

H

H

H

H

ii

■i

M

II

"s deck plate

-1

-1 -1 -1 -1 -1 -1 -1 — 1

Ao.d

(HPa)

150 170 200 200 240 240 300 300 340

Aos

(Ml»a)

115

142 107 167 200 200 250 250 283

»1

3000000 947000

135000

245000

N 2

Cycles

470000

N 3

4570000 1432000

.1047000

464000

»A

>5700000

>5670000 5031000 1676000 1069000 1211000

591600 560000 527000

II

II

II

II

II

II

II

II

II

II

U.Lg.

U.Lg.

U.Lg. U.Lg.

0,1 0,1

_ i

-1 -1 -1 -1 -1 -1 ™ x

0 0

amp.var. var.amp.

100 240

200 240 240 300 300 300 340 340

215 199

150 180

167 200 200 250 250 250 283 203

198 105 227 255

750000 22100

850000 208000 403000

248000 345000 135000 120000

522000

950400 66000

4007000 1200000 1900000

712000 434700 410000 405000

1065600 232000

4293000 1562000 2170000

440000 796000 447900 517000 470750

1678000 1315000

753000 791000

TABLE 3.3.a. Failure in the deck. (Ao . at point A ) .



Lab.

IRSID

U.Lg.

0)

U.Lg.

U.Lg.

RS(weld root)

0,1

- 0,63

- 0,71

- 0,96

- 1,0- 0,57

- 0,61

- 0,58

- 0,34

0

0

0

Var.amp.

Var.amp.

Aod(MPa)

240

296

302

333

150139

128

144

282

157

224

206

Aos(MPa)

180

240

245

280

152144

102

142

225

136

182

177

228

252

N, N2Cycles

282.000

2.100.0001.850.000

5.600.000

N,

1.230.400

290.000

276.000

400.000

7.600.000>14.750.000

>18.000.000

1.610.000

510.000

8.900.000

610.000

1.315.000

• 465.000

465.000

TABLE 3.3.b. Failure at the weld root(Åo at point D) ,



T A B L E 3 . 4 .

Fa t igue l i f e

C o n n e c t i o n S t i f f e n e r - D e c k p l a t e

T h i c k n e s s ( m m )

Plate

12

12

13

13

14

14

14

14

S t i f f e n e r

6

6

6

6

6

6

7

7

s u r f a c i n g

0

60

0

60

0

60

0

60

N u m b e r o f l o r r i e s ( 1 0 6 )

à°s ,

2a

71

44

199"

89

494

62

297

4 a d

20

75

45

219

103

647

73

393

27



R

"'M

D

-1

r

-1

-1

-1

0.1

-1

-1

0*(MPa)

150

17Ò

200

240

240

300

340

Initial

dei seta0

(mm)

0.050.100.150.200.40

0.050.100.150.200.40

0.050.100.150.200.40

0.050.100.150.200.40

0.050.100.150.200.40

0.050.100.150.200.40

0.050.100.150.200.40

NUMBER OF CYCLES AT FAILURE

F.M model

Propagation

7286000365000023880001766000908200

5031000255500016720001236000635500

316600Ó16080001052000777900399900

1884000957000626000463000238000

262000133000870006500033000

998000507000332000245000125900

69800035500023200017200088100

IRSID tests

Total

>5700000

>5700000

583100016760004293000

1069000121100015620002170000

232000

591600560000796000

470750

Propagation

20310007290003443000

107600012740001767000

209000

315000548000

350000

Tableau 3. 5 : Comparison of the tested and computed lifetimes

28 -



CLASS

CECM

3 6

4 04 55 05 66 37 18 09 0

AKs

( M P a v ^ T )

1 1 .1271 . 2 5 71 . 4 3 01 . 6 0 41 . 7 7 71 . 9 9 42 . 2 5 4

2 . 5 5 7

2 . 8 6 1

MONTLHERY

M i n e r

6

81 42 13 35 5

1 0 31 9 94 0 6

M.R.

3 03 13 43 84 45 57 8

1 5 9to

1 AK1 ^

( M P a v n T )

1 . 1 1 71 . 2 4 6

1 . 4 1 71 . 5 8 91 . 7 6 11 . 9 7 62 . 2 3 3

2 . 5 3 4

2 . 8 3 5

CARONTE

M i n e r

3 75 79 9

1 7 23 1 56 0 0

1 3 9 32 8 3 58 0 5 8

M.R.

1.882 1 82 6 9

3 7 1fa >wub>

W

Table 3.6 : Computed lifetimes by two models for various

S-N cu rves and Ak (bridg es of Montlhéry and Caron te)

a0

(mm)

0 . 1

0 . 2

0 . 3

0 . 4

0 . 5

0 . 6

0 . 70 . 8

0 . 9

K O

MONTHLERY

AK

1.430

>>

88

3 4

2 2

17

14

1 211

10

9

, , ( M P a / -

c eu 1 1

1 . 6 0 4

>>

1 1 8

3 8

2 3

1 7

1 4

1 21 1

1 0

9

"S )

2.254

>>

>>

7 8

3 3

21

16

141 2

11

9

CARONTE

AK (MPavnïï)s e u i l

1 . 5 8 9

1>>

>>

3 7 1

1 6 6

1 0 9

8 3

6 9

5 9

5 3

4 8

2 . 2 3 3

>>

>>

>>

4 3 6

1 8 7

1 2 5

9 6

7 9

6 8

6 0

Ta ble 3 .7 : Life tim es computed for various a and AKo s

- 29 -



I n i t i a lc r a c ka (mm)

0

0 . 1

0 . 20 . 30 . 40 . 5

0 . 60 . 70 . 80 . 91 .

L i f e t i m e s ( y e a r s )

T r a f f i c C hoi s y

3 6 2

1 6 51 0 4

7 66 1

5 14 54 03 6

T r a f f i c A n g e r s

8 2

3 82 41613

11988

3 3 7

Ta ble 3 .8 Lif etim es computed (AK = 0) fo r variou s a (Choisy).t o

CLASS

CECM

3 64 04 55 0

5 6

6 37 18 09 0

AKi

MPav'm

1 . 1 2 71 . 2 5 71 . 4 3 01 . 6 0 4

1 . 7 7 7

1 . 9 9 42 . 2 5 4

2 . 5 5 7

2 . 8 6 1

T r a f f i c

M i n e r

2 63 76 09 2

1 3 8

2 2 44 0 17 6 8

1 4 6 9

CHOISY

M . R .

1 0 71 1 21 2 11 3 2

1 4 7

1 7 62 3 43 6 17 5 4

Traff ic ANGERS

M i n e r

68

132 13 1

5 19 4

1 8 73 6 8

M . R .

2 42 62 83 03 4

4 25 9

1 1 53 3 3

Table 3.9 : Comparison of both models (Miner and F.M., Choisy)

30



Ch anf re in : sans

A n gle f i l : 6 0 'Te nsio n ce (V) : 29In te ns i t é (A) : 550Avance (cm/mn) : 50Energie (kJ/cm) : 19,1

F igu re 3 .1 : Ma crog ra phy o f we ldm en ts f o r op t im izedwe ld ing cond i t i ons

d ec k p l a te wh e e l a c t i o n

' s t i f f e n e r

JBBHB

wh e e l a c t i o n

F igu re 3 .2 : Behav iou r o f c r th o î r o p i e s te e l deck

- 31



- .Y l y

L i

ax e

c

a01x

¥ 'axe 1r-

point where stressesare calculated

^ - x

OX VIEW

r r f

L-kl

thickness :12mmgqqg f \J \ f \J VJ L th'ckness =6mn,

7 955 109 95.5

2100

Figure 3.3a : St ruc tu re used for the calcu la t ion

| (N/mm)

NJSls

Auoet real struct ure

/ / Icalcu la t ion s t r uct ur e

Figure 3.4 : 0 inf luence l ine 0 for a 10kNwheel c irculat ing on axis 2

Figure 3.3b : Detai l

3 2 -



Figure 3.5 :0 A Influence lines fop a wheel laxlsl ; y=0)

transverse repart i t ion

I

coco

Figure 3.6 :o 0-lnfluenc e lines fop a wheel lax lsl ; y=0) it ransverse repart i t ion

► X



n i

Ini

0 , 3 -A o m

s 9 8 N / m m

-—EL =0,115In i

0 , 2 -

dornage distribution (m = 3)

Figure 3.7 : S tr e s s Range Histog ram s a t P oint D'(var iable ampl i tude test)

34 -



Figure 3.8 : Test specimen and loading

poste 5

-Stress evolution in the deck plate : point A

-Stress evolution in th e trough : point D

Od 1 ("Pa)

poste 4

Fs lOkN

i i i i i i 0 1 Í » U Ö 15 » d (mm;

C U ("Pa) 400_

_ 9 0 .

_ 7 0

- 1 —

£5

paste 3

F.-IOkN

► 50 d (mm)

Figure 3.9 : Strain gauging of the specimen and nominal stress extrapolation

35



co

Aod

400

320

— 280CU

CL 240

200

i

LUCD

ŒCC

COCOLUaz

16 0

mo

12 0

10 0

80

70

CO 60

50

"1 I I I I N U 1 I l I I 1111 I I I I I 1111 1—I I I I I l i

40 < 1—1 I I I 1111y 2 3 456 8

Ê 0 ULg R=0 Ao CteL ® ULg Ao Var.: + IRSID R=-1: © IRSID R=0.1

105 2 3 4 5 6 8 R

i o5

i o6

i i i im2 3 4 5 6 8 7 2 3 4 5 6 8 n

i o7

I O8

CYCLES NFigure 3.10 : Crack initiation at the w eld toe : A o d



co

A os

4 00

3 2 0

-— 280CU

CL 240

2 0 0 h

LULD2 :CECO

COCOLUo c

160

140

120

100

80

70

cn

40

60 -

50 -

10

CYCLES NFigure 3.11 : Crack in i t ia t io n at the weld r oo t : Ao<



s> S ş

Ve^

Figure 3.12 : Inf luence s u r f a c e o f th e t r a n s v e r s a l e s t r e s s along th e s t i f f e n e r - p l a t e connection

38



crack length [m] t raff ic : Choisy

0 . 0 05

0 . 0 04

0 . 0 03

0 . 0 02

0 . 0 0 1

;

—

- -

"

-

'

— K

— K

- - K

- 0 .

- 1 . 1 2 7

- 1 . 6 0 4 - 2.254

r i ■ ■

l , / '

/ ,' -'

' /

/ ' i / ' /

/ ' / / / !

/ / '

/ 1 ' / /

/

, , . , ' .

" ' ' ' ■ i - - — -

;

;

/

," . ■"

•

;

,

;

■

'

-

0 100

Crack propagat ion depending on the A K S

20 0 , , 3 0 0 time [years]

crack length [m]

0 . 0 0 5 t—

A K S = 1 . 6 0 4

0 . 0 04

0 . 0 03 •

0 . 0 02

0 . 0 0 1

Choisy Angers

J /

' 0 50 100 time [years] 150

Crack propagat ion f o r the two t r a f f i c s

Figure 3.13 : Evolut ion of the crack l eng th with the time lo r the cycles number)

- 39 -



1». CONNECTION STIFFENER-STIFFENER.

H.1. Types of connections.

In most large orthotropic steel bridge decks field splices are

necessary, because transportation of the complete bridge from the shop to

the site is seldom possible. In general, both longitudinal and transverse

field splice have to be made. In the longitudinal splices only the deck

plate has to be connected. This is mostly done by butt welding and as the

weld is accessible both from above and from below, a good quality weld can

be achieved. The same applies to the transverse butt splice weld in the

deck plate. However, at the transverse splice, the longitudinal ribs have

to be connected as well (figure il.1). With closed ribs, which are usually

used in modern bridges, the most appropriate way of splicing is by welding,but as the welds can only be made from the outside in an unfavourable

overhead position, the quality of those welds will be dubious. Depending on

the location of the splice in the deck, the load on the splice can have a

fluctuating part due to the traffic load, dominating the static loading, so

conditions for fatigue damage are present. Stiffener splices occur

frequently in a bridge deck and fatigue cracks in this connection have

recently been found. An investigation into the fatigue behaviour of

stiffener splices is therefore required.

The Dutch and the Italian partners investigated the fatigue behaviour of

the field splices in this ECSC research. The Italians studied a triangular

shape of the trough ; and trapezoidal shaped troughs were tested by the

Dutch. A triangular shape has been studied by Cunninghame in 1982 [2 2 ].

In 1988 a Japanese IIW document was published which contained fatigue

results on trapezoidal ribs [ 2 3] . Test specimens and fatigue results of the

ECSC research are compared and presented by plotting S-N curves. In these

graphs the Eurocode fatigue design curves are used as a reference.

M.2. Stress determination.

4.2.1. Measurements on a Deck Panel in Pisa.

The specimen tested was a full size portion of an orthotropic

bridge deck (see fig. 4. 2 ). The steel plate, 2000 mm wide, was stiffened

- 40 -



longitudinally by three torsionally stiff ribs with triangular cross

section, spaced at 600 mm centres, and by two flat stiff eners (200 x 12

mm), to reproduce the transverse continuity of the deck.

The stiffeners were made of 6 mm. plate which had been cold formed. These

passed through the cross beams to which they were welded with fillet welds.The cross beams were a double »T' section laid at a centre distance of 3500

mm. The 12 mm thick web of the crossbeams had 25 mm radius circular

cut-outs at the apex of the troughs.

A 100 kN proof load was applied, using a hydraulic jack, through a

200 x 300 x 50 mm thick reinforced neoprene plate : the load was controlled

by a force ring gauge.

Forty five loading positions were tested on the deck ; strains were

measured at each location.

Strains were measured using electrical strain gauges located on

the deck plate, at the apex and on the webs of the stiffeners.

The deflections were measured at 88 points located under the stiffeners and

the cross-beam.

Diagrams and tables of the deflection and stress measurements are

presented in [J|].

1.2.2. Calculation of the stresses.

In order to study the static behaviour of the deck undergoing

examination, an extension of the classic HUber's method was adopted [18]

where the variable section of the crossbeams is taken into consideration,

as well as their shear strain. Using the simplifications introduced by

Pelikan and Esslinger [19], one obtains [20] the definition of a simple

analytical model to calculate the influence surfaces of the continuous

orthotropic deck on flexible cross beams. The differential equation that

governs the problem was resolved using the Levy's method [21]. The sections

of the theoretical deflection influence surfaces of the central ribs an the

stress influence surfaces at the apex of the central rib are presented in

- 41



The sample tested was also analysed using F.E.M. The code used was

MARC implemented on an IBM 3090 computer, whilst the finite element adopted

was a thin shell element with eight nodes with zero degrees of freedom

(element n°72 of the MARC library). The load (100 kN) was applied on a 20 x

30 cm. rectangular surface with its centre coinciding with that of thedeck. A comparison between the results obtained in this way and the

analytical and experimental ones, reveals a close correspondence. Thus, for

theoretical investigations, the analytical method was chosen, because of

its accuracy and its simplicity.

Some theoretical and experimental influence lines are compared in

figure 4.3. (deflections) and in figure 4.4. (stresses at the apex of the

central rib). In the diagrams, .the theoretical curves are shown as acontinuous line, while the experimental results are shown as a dotted line.

A study of the results reveals a close agreement between analytical and

measured values.

4.2.3. Stress spectra.

The stress spectra was obtained by means of the University of

Liège simulation program. The level crossing and the Rain-flow histograms

were calculated for the Rheden traffic on the influence line of the detailtested at T.U. Delft (Fig. 4.5. : bending moment at mid-span of a

continuous beam).

4.3. Test results of the University of Pisa.

4.3.1. Fatigue tests on type ' B' specimens.

4.3.1.1. Test specimens.

The Fe 51 OC specimens, 2000 mm long, are made of a triangular rib

obtained by the cold forming of a 6 mm thick steel plate welded to a top

plate 600 mm wide and 12 mm thick (figure 4.6).

Along the centre line of each specimen, a type I or a type II joint was

shop fabricated using the same procedures as those used on site.

In Figure 4.6 shows details at the tested joints the procedure. In type I

joints, the ribs stop about 100 mm short of the end of the deck plate.

42



In type I joints, the ribs stop about 100 mm short of the end of the deck

plate. The two ends of the top plate - one of which has the backing strip

welded to it (S1 w elding ) - are positioned w ith a gap of about 6 mm, and

then automatically welded.

The missing rib element is then inserted and manually welded in the

overhead position (backing strip weld S 3 ).

In type II joints, the top plate is about 100 mm short of the end of the

stiffener.

The stiffener webs are butt welded, with complete penetration manually

with coated electrodes. First, the internal part of the ribs is welded in

ascending vertical position. The root of the internal weld is ground

before the external weld is placed in the overhead position. The joint is

completed with the insertion of the missing top plate portion, theexecution of the S1 flat position welding and the manual overhead

remaining welding S2 between the top plate and the ribs.

All the w elds were checked by means of visu al and magnetic controls. The

butt weld, were also 100 % X-rayed and repaired where found (once only)

to be unacceptable according to the UNI 7278 Italian norms.

The fatigue tests at constant stress amplitude, were carried out on nine

type I joint specimens and on eight type II joint specimens.

The specimens were simply supported (at 2400 mm span) and the fatigue load

applied at two points 200 mm apart of the centre line using a pulsating

hydraulic jack w ith a frequency of 1 H z.

The load was applied though 50 x 60 mm. rectangular pats of neoprene, 10

mm. thick, between the mobile jack head and the top plate.

The strains at the apex of the ribs were measured using electrical strain

gauges, so placed as to determine the nominal stress amplitude without

taking into consideration the presence of local peaks of tension.

During the tests the minimum nominal stress was kept at a constant 1.5

kN/cm2 for all of the specimens.

Each test was stopped when failu re was reached, as recognized by the

specimen's loss of stiffness (an increase of one centimetre in the maximura

deflection under load), or when eight million cycles had been completed

without breaking.

4 .3.1.2. E xperimental R esults.

In table 1.1. the type of joint, the nominal stress range at the

lower part of the weld and the number of cycles to failure are shown foreach test.

43



In specimen 1 the first crack appeared in the fillet weld of the backing

strip, while a second crack appeared and propagates in the weld ; in

specimen 2, cracks initiated in both welds but only one crack propagated.

In specimens 3, 6, 7, 10, 11, 12, 13, 15, 16 and 17 cracks started in the

weld, at the apex of the rib, and propagated in the weld.

In specimens 4, 8, 14 cracks started at the apex of the rib in the parent

metal and propagated in the plate.

In specimen 5 the crack started in the weld at the apex of the rib and

propagates both in the weld and in the plate.

Specimen 9. no cracks occured in the rib, but three longitudinal cracks

appeared in the top plate : two at the position of the rib to top plate

connections and the third in the middle of the plate.

In figure 4.7 the results obtained and the mean life curves concerningtype I and type II joints are reported.

4.3.2. Fatigue tests on type 'A' specimens - (full size panel).

4.3.2.1. Test specimens.

In order to check the application of the fatigue results obtained

on type B specimens to the real deck joints, fatigue tests were carried

out on two large specimens with type I joints which were obtained by

cutting the specimen used in the static'test across at the centreline (seefig. 4.2).

The panel was supported on two crossbeam with an overhanging canteliver

section which was loaded by two concrete blocks of 54 kN total weight.

The pulsating fatigue 'load, applied through a 20 x 30 cm. rectangular

plate placed in the middle of the span, caused a nominal stress range at

the lower apex of the central rib weld equal to 22.5 kN/cm2 and a minimum

stress of 1.5 kN/cm2, reference being to the stress induced by the ballast

and the deal weight of the panel.

4.3.2.2. Experimental results.

The first specimen failed after 240000 cycles, the second one

after 260000 cycles. In both cases the crack initiated in the weld at the

apex of the rib and propagated in the weld toe.

In figure 4.7, the results obtained on type B specimens with type I joints

and those obtained on type A specimens are compared.

The comparison reveals close agreement, within normal experimental limits,

between the results obtained on the two types (A and B) of specimens.

The slightly lower fatigue life noticed on type A specimens is probably

due to the higher level of residual stresses present in these specimens.

- 44 -



4.4. Test results of the T.U. DELFT.

4.4.1. Tests specimen.

The specimens for the bending tests were single rib specimens as

depicted in figure 4.8. Two types of field splices were selected ;Type A : Butt splice with backing strips with varying root gap (0, 2 and

4 mm).

Type D : Butt splice with a V-groove and backing strips (root gap 4 m m ) .

The ribs were a rolled trapézoïdal section (steel grade Fe 510)

the dimensions of which conformed to F.K.H. - Trapezprofile nr.

2/325/6. As the usual spacing of the ribs is 600 mm, the width

of the deckplate in the specimens was also 600 mm, in order to

get the same position of the neutral axis as in an actual bridge

deck. During the fabrication of the test specimens, the welding

conditions on a real bridge deck were imitated because test

specimens made with special care under favourable conditions

would not be representive.

The ribs were welded on a 8400 mm. wide plate, at the proper

spacing of 600 mm., in a downhand position. Each rib was in two parts with

a gap in between to make the splice. Then this assembly was turned over

and the field splices were welded hampered by the adjacent ribs as it

would be on site. The welds in the bottom of the ribs were made in the

overhead position and the welds in the webs of the ribs were made by

upward welding. After completion of the splices, the assembly was cut into

fourteen test specimens and the end plates were fixed with fillet welds.

The test specimens were made by a fabricator with experience in making

orthotropic steel bridge decks.

4.4.2. Testing and measuring equipment.

As mentioned, a four point bending test was chosen to study

fatigue in the rib splices. Due to the fact that tests in the region of

ten million cycles were planned, two test rigs were buil t. Part of the

experiments, mostly the variable amplitude tests, were carried out using

servo hydraulic test equipment operating in closed loop control with load

feedback. The main part of the constant amplitude tests was executed with

loading equipment from Losenhauser. To avoid any secondary effects, all

supports in the test rigs were provided with hinges or roller bearings.

- 45



Each test specimen was instrumented with a number of strain gauges,

varying from 9 - 52 locations, before testing. Strain measurements were

carried out dynamically to check the applied stress range. Furthermore,

strains were monitored 2H hours a day to obtain information about and

location of crack initiation.

Measurements of crack growth were carried out during a number of

tests. This was done by periodic visual inspection with a magnifying

glass. It was only possible to measure crack length at the surface of the

specimens.

As far as possible four stages in fatigue failure are expressed in

number of cycles :

NI : Moment of crack initiation given by 10 % reduction in strain measured

by the gauge nearest to the crack.

N2 : Moment of visuable crack becomes visible.

N3 : A surface crack length of 50 mm.

N4 : End of test with extensive through cross section cracking (leading to

loss of specimen stiffness causing limitation of the actuator stroke)

and/or loss of symmetry (causing unacceptable side load on the

actuator bearing).

i|.i*.3- Constant amplitude tests.

i».11.3.1. Tests results.

The fatigue results of the constant amplitude tests are presented

in figure H.9. and table H.2. by plotting the a S-N relationship, on a

log-log scale. The normal stress in the bottom of the rib in the middle of

the splice, due to pure bending, was chosen as the main stress parameter

in this figure.

In all the specimens, the crack began in a weld between the rib

and the splice plate. In most cases, the starting point was located just

above the bend at the bottom of the rib. At this location the stresses are

about 20 % lower than in the middle of the bottom weld, where cracks

might be expected to start because of the higher bending stresses. In one

case the crack started in the side of the web at a location where the

stresses are about 70 % lower than in the middle of the bottom weld.

With some specimens it was possible to continue the testing forsome time to study the behaviour of the crack. The cracks propagated more

in the side weld than in the bottom weld, eventually the crack in the side

weld propagated into the base metal of the rib.- 46 -



To study the unexpected crack initiation points, some of the

cracked specimens were cut open. Inspection of the inside revealed local

defects, such as bad penetration and slag inclusion. The crack location

can be further explained by the welding procedure. The welds started are

finished just above the bend near the bottom of the rib and a tack weld

was also placed at this location.

Depending on the weld geometry and root gap of the weld, the first

visible crack N2 was noticed at 51 - 98 % of the number of cycles to

failure of the welded connection (N-U). The moment the invisible crack was

detected by measuring strains (N1) varied between 5 and 75 % of the number

of cycles failure (NU). For two specimens no cracks were detected by gauge

but one of the welds of these tests specimens failed very suddenly over 50

- 75 % of the total length of the weld. This phenomenon can be explained

by the poor quality of these particular welds.

1.1.3.2. Effect of the root gap.

Comparing the results of the fatigue tests from figure 1.9. it

appears that a weld with a zero root gap results in a fatigue life far

below the weld with a root gap of 1 mm. (a factor 18). A welded detail

with a root gap of 2 mm. gave variables results. At a level of 163 MPa it

gave a fatigue life comparable with the fatigue life of a detail with a 1mm. root gap but at a lower stress range level, about 110 MPa, the fatigue

life was a factor of 12 lower (Fig. 1.10).

1.1.3.3. Effect of the weld geometry.

Changing the weld geometry by using a V-groove did not give the

expected improvment. However, it is clear that it is easier for the welder

to make a better weld using a V-groove as a butt weld. The disadvantages

are that both sides of the detail have to be prepare and it is necessary

to use more welding material.

1.1.3.1. Fatigue limit.

From the results it appears that for the specimens with a root gap

of 1 mm. as well as those with a V-groove, no fatigue cracks were disco

vered at a level of about 90 MPa after testing for ten million cycles. It

can therefore be concluded that for this type of detailing we are almost

47 -



in the neigbourhood of the constant amplitude fatigue limit. For specimens

with zero root gaps however, several cracks were discovered in the bottom

side of the rib at a very early stage at a stress range level of 83 MPa.

Due to the bad results for this detail it is expected that there will be

no fatigue limit.

4.4.3.5. Comparison with previous Dutch tests.

In figure 4.10. the test results of the specimens with a root gap

of 4 mm. are compared with previous Dutch tests [24]. It appears that the

fatigue failures of the specimens tested at the high stress range levels

fall between the scatter band of the previous tests. It can therefore be

concluded that these fatigue results can be used for a weld classification

for this detail.

4.-4.3.6. Weld classification : S-N curves.

Ignoring the fact that cracks did not initiate at the bottom side

of the rib, it can be concluded that for the rib splice with a root gap of

4 mm., a weld class 80 (according to the Eurocode 3) can be recommended.

If the crack initiation point is taken into account the classification

reduces to a weld class 63. This values are to compare with the values

given in last draft of the Eurocode.

4.4.3.7. Calculated fatigue life.

Using the simulated stress spectrum for the field welded splices

(fig.4.5) and assuming that the detail can be classified as class 80

according to Eurocode 3, a fatigue life of about 75 years can be

calculated. Further optimisation of the detail seems to be unnecessary.

However in a lot of existing bridges welds were made with small root gaps.

Here the fatigue lives will be very short ; first failures have already

been discovered. To assist with the maintenance of these bridges it is

necessary to know how to repair those welds in an economic way and how to

calculate the remaining fatigue life after repair.

4.4.4. Variable amplitude tests.

4.4.4.1. Load spectra.

In the working group meetings it was decided to use load spectra

from the earlier ECSC Phase 1 and 2 research for the variable amplitude

tests of the third Phase. In the computer assessment programme of theUniversity of Liege, it was decided to use the traffic flow of the Rheden

48 -



Bridge. Results of the simulation for the field splice in a rib are

given in section 4.2.3. Analysis of the spectrum (figure 4.5) showed that

the stress range < 18 MPa caused only 7 ? of the total damage of the

sepctrum.

However, these stress range classes amount, to 84 % of the total number of

cycles. So leaving out these classes saves a lot of testing time. The

remaining stress ranges (18 MPa - 82 MPa) had to be raised a level above

the assummed constant amplitude fatigue limit. For testing techniques it

was necessary to leave out the highest stress range classes, which

accounted for only 5 J of the total damage. As well as , the simulated load

spectra, for one of the specimens a measured spectrum was used. A review

of all sepctra used is given in table 4.3.

To compare the variable amplitude tests results with the constant

amplitude ones, the applied stress spectra were analysed'according to

Miner's rule in the two different ways explained in section 3.2.3«

4.4.4.2. Tests results.

The fatigue results of the variable amplitude tests are presented

in figure 4.11. by plotting the S-N relationship, on a log-log scale.

The location where the cracks initiated and the crack development

were the same as those found in the constant amplitude te sts . The first

visible crack N2 was noticed in all cases at about 80 % of the number of

cycles to failure of the welded connection (N 4) . The moment the

(invisible) crack was detected by measuring strains (N 1) , varied between

69 and 76 % of N4.

4.4.4.3. Comparison with C.A. - tests.

- Test specimens with a root gap of 4 mm.•under simulated spectra.

There is a good agreement between the constant amplitude tests and the

variable amplitude tests using the Miner's calculation. Furthermore no

cracks were found in a specimen A.1.8. loaded to twice the simulated

spectrum after 48 million cycles. The maximum stress range of the

spectrum of this specimen was 132 MPa.

- Test specimen with a root gap of 4 mm under measured spectra.

Specimen A.1.6. was tested with the measured spectrum of the Forth

Bridge. Comparing the fatigue strength with the constant amplitude test

results, the spectrum showed to some degree a better fatigue life.

- 49



* Test specimen with V-groove under simulated spectra.

The fatigue endurance of this specimen D1.1.-V w as about 8 times that

calculated from the constant amplitude test on this type of weld

geometry using Miner. Comparing the fatigue behaviour of the relevant

speoimens, the fatigue behaviour of the V-groove specimen is almost H

times that of the specimen without the v-groove. More tests on this

detail are necessary to confirm this difference.

Comparing the results it appears that applying the measured spec

trum (A.1,6) resulted a very small difference. Using the measured

Spectrum (specimen A.1.6) instead of the simulated spectrum gives not a

significant difference in fatigue life.

i\.i\.5. Conclusions.

Cracks in these welds always started on the inside where

inspection is impossible. When a visible crack is located the total weld

will fail very soon. Repairing the weld must therefore be done as soon as

possible.

- Constant amplitude test ;

The results of the tests showed that the root gap of these kind of

welded connections must have a minimum width of 4 mm. to achieve a

welded detail class 8 0 according to Eurocode 3. If a weld can be made

without penetration defects and slag inclusions the classification may

be higher. F or the test specimens with a root g ap of 4 mm., no cracks

were found at a stress range of 9 0 M P a , (compared with the 6 0 M P a

constant amplitude fatigue limit of class 8 0 ). S o a knee-point at tw o

million cycles instead of five million cycles is possible. Changing the

weld geometry did not give the expected improvement, however the number

of tests (two) is too small to draw a definite conclusion.

" Variable amplitude tests t

F or the traditionally welded connection with a root g ap of H mm., there

is good aggreement between the constant amplitude and variable amplitude

tests, using Miner's calculation. I n this part of the programme the

V-groove (one test) resulted in a much better result than the welded

connection tested under a constant amplitude load. More tests are

necessary to confirm this difference. The difference in fatigue life

using a simulated spectrum seemed to be small.

50 -



H.5. Comparison with other research programs.

In addition to the ECSC research described in this chapter, the

fatigue behaviour of field splices in ribs of orthotropic steel decks

has been studied by Cunninghame [22] and Yamada [2 3]. A triangular shape

was studied by Cunninghame in 1982 and in 1988 a Japanese IIW document,

published by Yamada, contained fatigue results for trapezoidal ribs. The

results of the tests are given in figures 4.12 to 4.14..

From the UK-research it can be seen that, with the exception of

one test, all results are situated above the class 125 of the Eurocode

S-N curve. As with the Dutch results, the location of the fatigue cracks

was unexpected. All the failures initiated at the root of the weld on

the flat web of the stiff ener rather than at the more highly stressed

apex. Examination of several fracture surfaces showed no single

initiation point. It is concluded that the crack initiation point is

determined by residual stresses due to the welding procedure. Gathering

the "Pisa-tests" (type I) and the "UK-tests" together it can be

concluded that for triangular shaped ribs containing a weld with backing

strip a Eurocode class 112 can be considered (figure 4.12).

Confirming the Dutch research, the great influence of the size of root

gap on the fatigue strength was found in the Japanese research.

Furthermore the fatigue strength was affected by the residual stresses

in the direction of the rib which depends largely on the welding

sequence. Comparing the results with the Eurocode S-N curves, it can be

concluded that field splices with a root gap of 3 mm. or more are above

the class 71 curve. Gathering the "Delft-tests" and the "Japan-tests"

together it can be concluded that for trapezoidal shaped ribs containing

a weld with a backing strip and a root gap greater than 3 mm., a

Eurocode class 71 can be considered (figure 4.13). However with small

root gaps the classification can drop to 36 or less (figure 4.14).

4.6. Conclusions.

A lot of data are available concerning the field welded rib

joints. A first comparison showed ;

- The butt weld connection gives lower fatigue behaviour as the connec

tion with braking strips (fig. 4.7) ; but the fatigue behaviour of

this last connection seems influenced by local defects, that are

depending of difficulties of realization in a bridge working site.

- 51



The better behaviour of the triangular ribs compared with the

trapezoidal ones, for a connection with bracking strip and a root gap

greather than 3 mm. :

¿ 0 »11 2 N/mm

2

for triangluar ribs ;cÄ a = 71 N/m m 2 for trapezoidal ribs (that corresponds to

cEurocode 3 prescription).

- The reduction of the fatigue strength when the root gap is smaller than

3 mm. ; La < 36 N/ mm 2.

- The influence of residual stresses and defects due to the welding

procedure.

- Using Miner's calculation there seemed to be a good aggreement between

the results of the constant and variable amplitude tests.

More detailed analyses of the available experimented data are

needed to give a general conclusion for design and practical recommenda

tions.

From the above, it follows that for the orthotropic deck details

whose dimensional characteristics are described, it would be desirable to

define in the codes, S-N curves which take into account the type of joint,

the presence of defects and here control and, above all, the construction

procedures and thus the level of residual stress (for example : the

starting point of welding). The S-N curves should be lower for those

details which do not satisfy minimum quality requirements. Naturally,

because of the complexity of the problem, the conclusions proposed here do

not pretend to be in any way definitive, but rather a reference point for

future investigation.

52 -



Speci-.en

Jí.

1

->

3

4

J o i nz

Type

I

I

II

II

5 ! I] C

5

7

S

3

10

11

» 113

14

15

16

17

I T

I

o-rsnçeI KW/en = ]

19.50

22.50

22.50

19.50

17.50

17.50

22 . 50

II 1 14.00

I

T T

I

Ţ

II

1 1

I

II

I

17.50

19.50

19.50

Cycles ac failure

13S0000

53S0OO i

251000

939700

321000

459500

463000

6040000

8100000

335000

1200000 "

22.50 1 S560CO

22.50

17.50

19.50

305000

1930000

1657000

12.50 1 7073000

19.50 1 1622000

Test results o f Pisa

TABLE 4 . 1 .

- 53



SPECIMEN

A.1.1.

A.1.1.

A.1.2.

A.1.2.

A.1.3.

A.1.3.

A.1.4.

A.1.4.

A.2.1.

A.2.1.

A.2.2.

A.2.2.

A.3.1.

A.3.1.

A.3.1.

A.3.2.

A.3.2.

A.3.2.

A.3.2.

D.l.l.

D.l.l.

D.I.2.

D.I.2.

VELD

1.

2.

1.

2.

1.

2.

1.

2-

J

2.

1.

i »

1.

1.

1.

2.

1.

nl

0.157

2- *

V•

-

9.085

•

-

-

0.180

0.239

.

-

-

0.713

0.035

0.035

-

0.175

0.180

0.203

0.223

-

0.728

-

n3-

•

9.085

-

•

-

0.180

0.244

.

-

-

-

0.039

0.029

•

0.180 .

-

0.253

0.240

-

-

-

n4> 0.833

0.833

9.298

> 9.298

>29.582

>29.582

0.245

> 0.245

> 0.668

0.668

> 0.779

0.779

0.445

0.445

> 0.465

0.327

0.327

0.327

> 0.327

>10.990 .

>10.990

0.798

> 0.798

At>r[KPa]

150

Ï50

105

105

90

90

233

233

163

163

110

110

153

153

153

S3

83

83

83

lêo

CA?[ 3]

'M'Vi' l ìî4 :u

4 ƒ

4 J

2

2

2

2

0

0

0

0

0

0

0

H '

1 J

to - nominal stress ranga bottoaside trough.

n - number of cycles x 10 .

n. - initiation by strain gauge.

n, - first visual crack.

n, - length of crack equal to 50 va.

n, - failure of the veld.4

Table 4.2. Results constant amplitude tests

5 4 -



STVÜLATE3 S7ECTP.A; TABLE 4 . 3 .VtRTABLE AMPLITUDE TESTS ! T- U. £1 ITT

N U M S E S

O ?

CYCLES

E1 3 0

s a i a

5 5 7 0

5 41 0

51 2 0

« 6 2 0

3 5 0 03 7 0 0

3 3 3 0

2 P O O

2 5 3 0

1970

1 6 7 0

1 5 5 0

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D E C K P L A T E

L O N G IT U D I N A L R I B

L A T E R A L R I B

R E L D W E L D E D J O I N TM A I N G R I D E R

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D E C K P L A T E

B A C K I N G S T R I P

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W E L D J O I N T

L O N G I T U D I N A L' R O U G H R I B

C O N N E C T I N GR I B

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56



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57



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r e s u l t s ( d e f l e c t i o n s )

- 58



0

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59



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I I I I I I I I I REGRESSION LINES

CONSTANT AMPLITUDE TESTS

T I I I I I I I I 1 I I I I I I -

J U - D E L F T S t e v i n L A B O R A T O R Y S t e e l S T R U C T U R E S

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Figure 4.11 : Field-welded rib joints of orthotroplc stee l bridge decks

3 4 5 6 7 9 R

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[ N ] — - >

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F I E L D - W E L D E D RIB J O I N T S OF O R T H O T R O P I C S T E E L B R I D G E D E C K S

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,14 : Field-welded r ib Joints of or thotropic steel br idge decks

2 3 4 5 6 7 9 o

1 0

[ N ] - — >STST-002.GRA



5. CONNECTION STIFFENE R-C ROSSBEA M.

5.1. Types of connection.

Bridges with orthotropic steel decks buil t in Europe used trapezoidal

or 'V' shaped longitudinal st iffeners. Sometimes in the past these longitudinal

st iff en ers were butted up to the transv erse s t iffen ers. Fatigue fai lure of these

troug h to crossbeam welds occu rred in heavily traffick ed bridges in less than

20 years. In other cases, the longitudinal st iffeners passed through cut-outs in

the crossbea ms and the conn ections were no long er w eakened by load carrying

welds at" each side of the crossb eam . Unti l now, no fai lures of the st iffener

through crossbeam connections have been reported, however, information about

the fat igue classificat ion of different types is unknown.

Tests at TRRL were intended to establish the fat igue behaviour of

thr ee connections typica l of the la te r typ es. They are i l lustr ated in Fig. 5.1 .

Recent research [26] suggests an improved form of the cut -outs in the

crossb eam , see type ' R' , Fig. 5.2. Te sts a t D elft were intend ed to get an

agreement of stress distributions and fat igue behaviour of three types of

connections (Fig. 5.2) and not to define a fatigue design curve. The work of

LBF was to study stress fields in the crossbeam in different shapes of

cut-outs in order to reduce stress concentrat ions (Fig. 3.3 and 3.1).

5.2. Stress determination.

5.2.1. Measurements on a deck panel at TRRL.

The three types of connection shown in Fig. 5.1. were incorporated

in the ce nt ra l crossbeam of a 15.2 m long by 3.1 m wide deck paneL Strain

gauges were instal led around each connection, on the web of the trough and

on th e crossb eam , 15 mm. from th e ro ot of th e weld. The main gaug es

instal led around co nnections 'A ' and ' B' are shown in Fig. 5.1 .Stat ic loads were applied to the panel through a single wheel and the

influence surface of stress was obtained for each gauge posit ion.

The panel was unsurfaced.

70 -



By superposition, longitudinal influence lines of stress were calculated

for the vehicle types described in the British Standard code of practice

for fatigue, BS 5*100 part 10 [11], Using the Rain-flow cycle counting

method, histograms of stress ranges were calculated for one million

vehicles distributed across the deck (BSS^OO, clause C.1.*0 with the

centreline of the distribution directly over the trough. The stress spectra

for the important gauge positions around connections 'A' and ' B' are given

in Table 5.1 •

5.2.2. Measurements on a UK bridge.

Strain gauges were installed around a trough to crossbeam connec

tion on a heavily trafficked UK bridge. The instrumented trough was located

under a slow lane wheel track. Gauges were installed at the positions ofhigh stress indicated in the panel tests (see Fig. 5.1).

Strains were recorded continuously for a two week period during which the

asphalt temperature ranged from -1.6°C to 18.0°C (mean 7.6°C). Stress

histograms for the 8 gauges are given in Table 5.2.

5.2.3. Local stress calculations and measurements in the web of the

crossbeam at LBF.

The stress level in the web near to the cut-outs is proportionalto the shear forces in the crossbeam and is practically independent of

bending moment. Additionally, these stresses are influenced by the loads

introduced locally into the crossbeam. Therefore the fatigue problem at the

two critical points (notches in the cut-outs and end of weld between the

stiffener and the web near to the cut-outs) may be investigated by using a

simply supported crossbeam together with short pieces of the stiffeners and

of the deck plate as a test specimen and as a model for calculation (Fig.

5.15). This crossbeam model may be loaded by definited shear forces to

simulate stress distributions and stress time histories similar to those in

bridges under real load conditions.

Two shapes of cut-outs in the web of crossbeams were investigated,

one which is commonly used (type II) and the other, an "improved" design

proposed by Haibach [26] after investigating crossbeams of railway bridges

(type I ) . The stresses due to shear forces in the beams at both types of

cut-outs were computed and measured, Fig. 5.3, 5.4, 5.5.

As a result the stress distribution around the cut-outs was found to be

antimetrie with respect to their plans of symmetry. The level of maximum

stress at the notches of both cut-outs is nearly the same but the volume of

highly.

- 71 -



stressed material is much larger in the conventional cut-o uts. At the end

of the welds (the other critical point of the connections) the new shape

produces significantly lower stresses than the conventional shape.

5.2.^. Stress Spectra.

The stress spectra were obtained using the simulation program of

the University of Lièg e. This program calculates the level crossing and the

Rain-flow histograms.

Two influence lines are tested with the Rheden traffic :

* influence line concerning stresses due to bending moment at the

stiffener on the support of a continuous beam. This has been

measured by the T.R.R.L (Fig. 5.6. and table 5. 5) .

* influence line of the support reaction of a continuous beam. This

has been given by the L.B.F. (Fig. 5.7.).

5.3. Test results of TRRL.


Full-scale test specimens were manufactured comprising a 1500 mm

length of deck plate, a single trough and a central crossbeam. Two types of

specimen were made with the detailing of the trough to crossbeam connection

representing the 'A' and ' B' connections of Fig. 5.1. Six millimetre,

single pass, Manual Metal Arc Welds were used for the trough to crossbeam

connection. Inspection of the welds showed them to be representative of

those that could be found on a typical bridg e. The specimens were loaded in

a reaction frame test rig, illustrated in Fig. 5.8.

Strain gauges were installed at the high stress locations around

the weld (and 15 mm from the wel root) in identical positions to the gauges

on the test panel. The loading on the specimens was arranged to produce a

similar distribution of stress around the connection as that determined

from the tests on the deck panel with the wheel load in the most damaging

position.

- 72



Five type 'A' and six type 'B' specimens were tested at constant

amplitude. A further four type ' B' specimens were tested at variable

amplitude. In all cases a 25 mm crack was defined as failure of that part

of the specimen.

5.3-2. Test results - type 'A' connection.

Fatigue cracks developed at four locations (see Fig. 5.9), though

not necessarily all in the same specimen :

crack a was a weld toe failure through the trough plate at the bottom of

the weld, initiating within 25 mm of the weld end. This crack was expected

from the high stresses found at this point in the panel tests.

Crack b was in the crossbeam plate at the bottom of the weld.

Crack c was in the trough plate at the top of the weld. It occurred in all

5 specimens, on both sides of the trough and on both sides of the cross

beam. In all cases cracks initiated at very low endurances. Stresses were

found to be much higher at this location than had been expected ; the

stresses were confirmed by measurements on the panel.

Crack d was in the crossbeam plate at the top of the weld. It occured in

only three specimens and at much longer endurances than crack c. It is

regarded as a secondary order crack.

For cracks a, b and c the endurances for a 25 mm crack are given in table

5.3 and plotted in Figure. 5.10, against the stress measured by the strain

gauge adjacent to the crack location. Estimated stresses at crack b were

used for four of the specimens. The data are compared with Eurocode S-N

curves. It is concluded that class 50 is appropriate for the failure in the

trough plate at the top of the weld (crack c) and class 125 for the weld

toe failure at the bottom of the weld (crack a ) . A high classification is

indicated for crack b from the estimated stresses at this point.

5.3.3. Test results - type '3' connection.

Out of six specimens, the two tested at the lowest stresses (95

and 100 N/mm2) were uncracked after 11,7 and 13,2 million cycles respecti

vely. The remaining four specimens all suffered weld toe failures through

the trough plate as expected (see Fig. 5.8.). Cracks initiated near the

apex of the trough. In one specimen, a second crack developed in the toe of

the weld at the crossbeam.

73



The endurance for a 25 mm crack is given in table 5.1 and is

plotted in Fig. 5.11. against the stress in the trough plate (at the apex

of the trough) adjacent to the weld. The results indicate a weld class 80.

Four specimens were tested under variable amplitude loading. The

applied stress spectrum was derived from the static load test data (section

5.2.1) for the gauge on the apex of the trough under the vehicle loading

from the Rheden Bridge (Section 5.2.1). The level of the spectrum was

raised to obtain failures in a reasonable timescale. The spectrum used in

the tests if given in Table 5.5.. It has an equivalent stress range La of

97.1 N/mm2.

1/3where Ao - [(1/Zn.) x E(n. a.')]

e i i i

The results of the variable amplitude tests are also shown in Fig. 5.11.»

plotted at the level of the equivalent stress range. The four results are

below the mean line of the constant amplitude tests the average endurance

suggesting that a fatigue life based on the Palmgren-Miner summation would

be optimistic by a factor of about 2. However, three out of the four

results are within the 95 % confidence limits for weld class 80.

5.3.1. Fatigue life calculations.

For each type of specimen, strain gauges were installed in identical positions on the deck panel and on the fatigue test specimens. It is

therefore possible, using the stress spectra calculated from the static

tests on the deck panel (Table 5.1.) and the weld classifications determi

ned from the constant amplitude fatigue tests, for corresponding gauges, to

calculate the fatigue lives of the connections for BS5100 traffic loading :

Type 'A' - crack a, bottom of weld through trough - > 120 years

- crack b, bottom of weld through crossbeam - > 120 years

- crack c, top of weld through trough - 5 years.

Type ' B' - crack through trough plate at apex of trough - 13 years.

The lives quoted are for a 2,3 % probability of failure and for

one million HGVs per annum. There is no influence from bridge deck

surfacing in these calculations.

Lives were also calculated using the stress spectrum obtained from

the measurements on the bridge. For crack c the life is calculated to be

280 years. This assumes that the traffic flow across the bridge and thetemperature of the bridge deck surfacing for the two week measurement

period is typical of that throughout the year. In fact the traffic flow is

- 74 -



k now n to be only about half a million HG V s and the su rfacing temperature

was below the annual mean.

Consequ ently, the connectio n fails to meet the U K design re qu irements

though failures in service are not expected to occur within the lifetime

(120 years) of the bridge.

The fatigu e life of 13 years calcu lated for the type '3 ' connectio n from

the constant amplitude tests would be reduced to around 6 years using the

average endurance data from the variable amplitude tests.

5.H. Tests results of the T.U. Delft.

Two series of tests were executed. In the first series the fatigue

load was about 50 % higher than in a real bridge. In all cases no cracks

w ere fou nd after testing for at least 12 milli ons cycle s. I n the second

seri es of tests the specime n of the first series w ere cu t in to tw o piec es

and were tested at a much higher level (at least H times the first loading

case). The testrig in that case was the same as used by the Transport and

Road Research Laboratory.

5.^.1. Test specimens.

The specimens for the bending tests of the first series were

single rib specimens. The ribs were rolled trapézoïdal sections of steelg rade F e 5 10 . The dimensions of the cross secti on w ere equal to a F .K .H . -

Trapezprofile nr. 2/325/6. As the usual spacing of the ribs is 600 mm., the

w idth of the deck plate in the specimens w as also 6 0 0 mm., in order to g et

the same position of the neutral axis as in an actual bridge deck. Three

types of specimen were made with the detailing of the trough to crossbeam

connection representing the ' S' , •T' and ' R' connections of figure 5.2.

These connections w ere manu ally w elded. B y cu tting the test specimen in

two pieces the specimens of the second series were made in the same rig as

used at TRR L (Fig. 5 .8 ).

The first test series were executed with loading equipment from

Losenhau ser. The second series were carried out u sing serv o hydraulic test

equipment operating in closed loop control with load feedback.

- 75



Each test specimen was instrumented with a number of strain gauges

before testing. Strain measurements were carried out dynamically for 2k

hours a day to obtain information about the time and location of crack

initiation.

Measurements of crack growth were carried out during a number of

the tests. This was done by periodic visual inspection with a magnifying

glass. It was only possible to measure crack length at the surface of the

specimens.

Where possible, four stages of fatigue failure are expressed

(see section 4.M.2) :

5.1.2. First series of tests.

As mentioned before no cracks were found in the first series of

tests at a load range of 50 kN. From these results it appears that the

highest stress is measured in the connections without cut-outs. The

stresses measured on the connections with the new design of the cut-outs

are a little bit higher than the design with the old cut-outs.

5.^.3. Second series of tests.

A review of the test results is given in Table 5.6. In this table

the following parameters are given :

- the fatigue load ;

- number of cycles for each stage of fatigue failure in the crossbeam and

the weld.

The results plotted on Fig. 5.13 and 5.1*4 correspond to N3.

The location of the cracks is shown in Fig. 5.12.

The cracks appeared in the crossbeam are influenced by the test specimen

and the testing, they are not considered her e.

The two specimens type ' S' suffered weld toe failures through the

trough plate as expected. Cracks initiated near the rounding of the

botton-web of the trough. Fig. 5.13 shows that the fatigue behaviour is

better than the highest class of the Eurocode curves.

- 76 -



In type 'T1 - connection, fatigue cracks developed in the

crossbeam in the rounding of the cut-out and the weld toe through the tough

plate at the botton of the weld. For the crack initiating at the weld, the

endurance is plotted against the stress measured at 12 mm from the weld toe

(Fig. 5.11).

Just as the type ' T' - connection, in type ' R' connection, fatigue

cracks developed in the crossbeam in the rounding of the cut-out and at the

weld toe through the trough plate at the bottom of the weld. For the crack

initiating at the weld, the endurance is plotted against the stress

measured at 12 mm from the weld toe (figure 5.11). It seems that also

fatigue behaviour of this detail is a little better as for connection T.

5.1.5 Conlusions.

Connection type S, without cut-outs has a better fatigue behaviour

as connections type T and R with cut-outs. The fatigue behaviour of

connections type T and R is very near. This conclusions are in opposite of

the results obtained at the V stiffeners (section 5.3) : type B, without

cut-outs gives a fatigue behaviour near at type T and R, but type A, with

cut outs at the apex, gives a higher fatigue behaviour as type B.

5.5. Test results of the L.B.F.

5.5.1. Test conditions.

Three test specimens were fabricated, nearly the same scale as the

orthotropic decks of real bridges, each consisting of a cross beam, 6

stiffeners, and a deck plate. One of the specimens had cut-outs of the

commonly used shape and two the improved shape . When the specimens are

loaded as shown in Fig. 5.15, four of the cut-outs are stressed at nearly

- 77 -



the same level. This is due to symmetry and the constant shear force.

Therefore up to four results can be expected from each test specimen,

provided cracks are early detected early and repaired.

Three fatigue tests were performed one with constant amplitudeloads applied to a specimen with the new shape of cut-outs and two using

identical load sequences of variable amplitude (derived from measured

traffic loads) applied to the other two specimens one with the old and one

with the new cu t-ou ts.

5.5.2. Test results.

Test results are given in Fig. 5.16, 5.17, 5.18 and 5.19 cracks

occured at both critical points of the connection specified earlier. Atthe end of the welds between the longitudinal stiffeners and the

crossbeam, cracks were observed only in the specimens with the new shape

of the cut-outs, althoug h this shape was developed especially to reduce

the stresses at this point. The stresses due to the external loads are

actually lower as shown by computation and measurement, Fig. 5.3, 5.4.

H ow ev er, it seems that the new shape increases residual stresses at the

end of the welds due to unequal heating during welding. High residual

stresses allow fatigue cracks to develop ever when the stresses from the

applied loads are low. With crack initiation, the residual stresses are

removed and the craks do not grow (this was observed during the tests).

In contrast the cracks initiated at the notches of the cut-outs

grew and would have destroyed the specimens if not repaired. These cracks

occured at the cut-outs of the new shape after 3-5 times more load cycles

than at the more commonly used cut-outs, which means a significant

improvement of fatigue life.

However high residual stresses globally distributed in all three specimens

were observed without any doubt. Their influence on fatigu e is so strong

that cracks occured at notches stressed only in compression due to the

external loads, while other notches remained crack free, although stressed

in tension at the same level. Therefore it is not clear, whether the

increase of life time is due to the new shape of the cut-outs alone or at

t least partially to a more favourable residual stress distribution.t

v . /78 -

i



5.6. Conclusions.

Three types of stiffener to crossbeam connection were tested at

TRRL, LBF and Delft : welded all round and with conventional and 'impro

ved' cut-outs in the crossbeam around the apex of the stiffener. Both

trapezoidal and 'V' shaped stiffeners were tested. In one of the research

programmes the main aim was to study the fatigue behaviour of the

crossbeam by applying loads to produce shear forces in the cross beam. In

the other two programmes the fatigue behaviour of the welds was studied by

applying bending stresses to the longitudinal stiffener.

Despite difference in the shape of the longitudinal stiffeners,

the weld sizes and probably different residual stresses due to different

weld procedures, a minimum weld classification of class 80 followingEurocode 3 can be considered for the welded all round connection. However,

a higher classification may be considered for connection type S

(trapézoïdal stiffener without cut outs) as for connection type A

(triangular stiffener with cut-outs at the apex). Conclusions for joints

in trapézoïdal stiffener are provisional, because only results for two

tests for each type of connection are available.

Most of the variable amplitude test on this type of connection

gave a good agreement with the constant amplitude tests applying Miner's

method of damage summation.

Because of the difference modes of failure from the difference

types of welded connections with cut-outs in the crossbeam it is not

possible at present to define an overall classification for this detail.

For specific details, classifications are given in previous sections where

possible.

The improved form of cut-outs did not result in a better fatiguebehaviour as had been expected. Cracks occured in unexpected locations in

the crossbeam connection in the new design despite low calculated stresses

at this point. Weld toe failures occurred at similar endurances in the

conventional design. For the failure in the crossbeam plate the new design

was an improvement over the conventional design but it was difficult to

manufacture this shape of cut-out without producing notches.

79



For the 'V* shaped stiffeners, the connection with cut-outs

behaved better than the fully welded type. This was not the case with the

trapezoidal stiffeners. A possible explanation is that residual stresses

may be higher at the apex of the 'V' stiffener than in the bottom of the

trapezoidal stiffener. Therefore a cut-out at the apex of the stiffener

which avoids residual stresses in this area gives a greater improvement in

fatigue behaviour for the 'V' stiffener.

To give a general classification for stiffener to crossbeam

connections it would be necessary to refer to nominal stresses rather than

stresses at specific points which cannot be easily calculated by the

designer. Further analysis of the results and additional research is

needed to give practical design 'recommendations.

- 80 -



09

STRESS

RANGE

°R(N/mm2)

0-10

10-20

20-30

30-40

40-50

50-60

60-70

70-80

80-90

90-100

NUMBER OF CYCLES

CONNECTION 'A'

GAUGE 13

6477245

481807

355667

255613

83593

22933

4063

0

0

0

CONNECTION 'A1

GAUGE 88

6118059

718470

279695

317424

175765

61490

38862

3899

0

0

CONNECTION 'A1

GAUGE 90

8456885

520645

288755

100389

0

0

0

0

0

0

CONNECTION 'B'

GAUGE 49

3119612

1254028

582721

302745

146634

205557

78214

21212

11404

1969

- For 1.000.000 HGVs

Centre of distribution of vehicles over centreline of trough

TABLE 5 . 1. S T R E S S E S F R O M T E S T S O N D E C K P A N E L



1

00r oi

STRESS

RANGE

°R(N/mm2)

8-12

12-16

16-20

20-24

24-28

28-32

32-36

36-40

40-44

44-48

48-52

52-56

56-60

1

15383

4614

1790

876

455

229

80

21

5

0

0

0

0

2

14680

10408

6254

3515

2116

1698

1082

490

177

42

4

1

1

3

20851

11301

5613

2236

992

325

34

3

0

0

2

0

0

NUMBER OF CYCLES

GAUGE NUMBER

4 5

17866

11785

3642

747

227

29

0

0

0

0

0

0

0

14454

5038

2071

1153

682

438

330

157

60

15

2

1

0

6

16098

10518

6064

3535

2569

1600

537

127

21

9

1

2

0

7

26099

12632

. 6787

2650

1463

664

186

28

2

1

0

0

0

8

20709

16227

10108

3341

936

518

120

17

0

0

0

0

0

TABLE 5 . 2 . S T R E S S E S F R O M M E A S U R E M E N T S O N B R I D G E



00

u

SPECIMENNUMBER

IA

2 A

3 A

5 A

8 A

CRACK a

NORTH SIDE

STRESSRANGE

O i

(N/mma)

12 5

I S O

2 0 0

14 4

16 0

CYCLES

TOFAILURE

X 1 0 d

>12.4

5 .2

1.8

1.4

6.3

SOUTH SIDE

STRESS

RANGEo=,(N/mmz)

11 3

14 2

19 2

17 5

15 1

CYCLES

TOFAILURE

X I O *

)12.4

> 6.3

> 2.5

2 .1

> 7.0

CRACK b

NORTH SIDE

STRESS

RANGE

(N/mm3)

1 3 7 -

1 6 4 *

2 1 9 -

1 5 8 -

17 6

CYCLES

TOFAILURE

X 1 0 &

>12.4

> 6.3

2 .3

2 .1

> 7.0

SOUTH SIDE

STRESS

RANGEOii»

(N/mmz)

1 2 4 -

1 5 0 -

2 1 0 *

1 9 2 *

17 7

CYCLES

TOFAILURE

xio<-

>12.4

5 .8

> 2.5

1.8

> 7.0

CRACK c

NORTH SIDE

STRESS

RANGEOT »

(N/ram=)

3 9 *

4 7

7 3

5 6

6 3

CYCLES

TOFAILURE

X I O *

6.3

3 .0

1.8

2 . 7 "

2 .9

SOUTH SIDE

STRESS

RANGEO n(N/mm*)

5 0 -

5 9

6 3

5 4

6 3

CYCLES

TOFAILURE

X I O *

8 .2

4 .0

2 .4

2 . 9 "

5 .2

denotes estimated stressdenotes extrapolated cycles

TADLE g-3 FATIGUE TEST RESULTS - TYPE 'A' SPECIMENS T.R.R.L



SPECIMENNUMBER

233343637B8B

STRESSRANGE

Ca(N/RWli*)

95150125115100200

CYCLESTO

FAILUREX10*

>11.700.902.801.67

>13.200.47

a) SPECIMENS TESTED AT CONSTANT AMPLITUDE

SPECIMENNUMBER

10312B13314B

MAXIMUMSTRESSRANGE

O i

(N/rem3)

245245245245

CYCLESTO

FAILUREXI0*

2.120.7S1.862.75

b) SPECIMENS TESTED AT VARIABLE AMPLITUDE

TABLE SA FATIGUE TEST RESULTS - TYPE -'B ' SPECIMENS TRRL

- 84



STRESSRANGEN/mro2

CLASSCENTREN / i r a n

2

CYCLESni

50-60

60-70

70-80

80-90

90-100

100-110110-120

120-130

130-140

140-150

150-160

160-170

170-180

180-190

190-200

200-210

210-220

220-230

230-140

240-250

55

65

75

85

95

105115

125

135

145

155

165

175

185195

205

215

225

235

245

23710

17440

12590

10920

9080

68505290

3440

3690

2340

"980

1410

570

710340

90

240

70

50

10

TABLE 5. 5 S T R E S S S P E C T R U M U S E D

I N T H E T R R L T E S T S "T&ft-l

- 85



TESTSPECIMEN

TYPE

v v y \ J

u

\j

\j

NR

SI

S2

R2

Rl

TI

T2

LOAD

[kN]

200

200

200

267

266

267 ■

LOCATION OF

CRACKS

VELD

CROSS BEAM

VELD

CROSS BEAM

WELD

CROSS BEAM

VELD

CROSS BEAM

VELD

CROSS BEAM

VELD

CROSS

BEAM

NORTH SOUTH

NORTH SOUTH

NORTH SOUTH

NORTH SOUTH

NORTH SOUTH

NORTH SOUTH BOTTOM

NORTH SOUTH

NORTH SOUTH

BOTTOM

NORTH SOUTH

NORTH SOUTH

NORTH SOUTH

NORTH

SOUTH

NUMBER OF CYCLES (xlOE06)

Nl

0.080

0.660

0.090

0.480

0.040

0.016 0.016

0.050

0.100

N2

1.658

0.809

1.157

1.125

0.271

0.771 0.174 0.653.

0.241

0.025

0.025

0.087

0.187

0.208

0.103

0.175

0.369

0.084

N3

1.732

0.879

1.184

1.157

0.474

1.190

0.711

0.250

0.564

0.087

0.112

0.348

0.606

0.397

0.351

0.676

0.376

N4

2.461

2.461

1.272

1.272

1.190 1.190

1.190

1.190

1.19P

0.572

0.572

0.572 0.572 0.572

0.700

0.700

0.700

0.700

0.856

0.856

0.856

0.856

Nl : Moment of crack initiation by 10 X strain fall of, measured in the

strain gauge nearest to the crack.

N2 : Moment of visuable crack initiation.

N3 : A crack indicating the number of cycles when a surface crack length

of ± 50 mm is reached.

N4 : End of extensive through cross section cracking.

Tablé S.'£.. Review of the second series results. T.U. Delft.

86 -



Type 'A'

Deck plate

00

Type 'C'

Type 'B'

Web of t roughWelded connect ion under test

and weld

Icul oul )

Figure 5 .1 : Types o f long i tud in a l / t ran sve rse s t i f f en er connect lon-TRRLAnd main gauges in test (specimen 'A'and 'B'and in the bridgestudied (specimen 'C'



00

co

and weld(cut out)

Figure 5.2 : Types of longitudinal / Tramsverse stlf fen er connection



00(O

load F = 200kNs. Fig. 5.21)

• measured at ~ 2 mmdistance to the edge

D measured on the edge

b^

calculated

Figure 5.3

COMPARISON OF COMPUTED AND MEASURED STRESSES AT THE CUT-OUTS OF THE NOW USUALLYUSED SHAPE (TYPE II)



o

%

c.

calculated ^

/S

/

CO I \

- c o m pression

\ \

X

xmeasured at - 2 mmdistance to the edge

load F = 200kNI s . Fig. 5.21 J

Figure 5.4

calculated

DD

\D

measured onthe edge

mea sured at ~ 2 mmdistance to the edge

COMPARISON OF COMPUTED AND MEASURED STRESSES AT THE CUT-OUTS OF THE SHAPEPROPOSED TO BE BETTER {TYPE I)



(O

measuredst ress

Figure 5.5

STRESS DISTRIBUTION IN THE NARROWST SECTION BETWEEN THE CUT-OUTS OF THE SHAPEPROPOSED TO BE BETTER



2 0 -

15 ■■

10 ■■

5 "

I 76.21

(152.AI

(226.6)

(30 4.81

___. influence line

/ V - " XN . used for —»

//' ^, the simulation./^' ' • ^ .

/ v f/r^X w (381.0) v \ a/ /

%

\ \

/

\ if

\\W/ m i /

. 1 /

-H 1 1—I—' I I i \ W / / l 1 I I h—I—I—I—I—I—I—h

:\':I/-K A1 M CO K7 75 3 7 11 15 19 25 \W/3S « 51 59

67 75 y ■■ fop a 20kN wheel W

3 -

"N/mm2 i n f luence l ines measu red at the T.R.R.L. wi th a 20kN whee l

Histogram Rain f low NI/NTOT

cu--

0.3 - •

0.2-■

0 . 1 - -

Rheden Traffic

' rT") n 11111111 N/mm H — » -

50 100 150 20 0 250

Fatigue damge , .D I

10 + 9

8 ■•

7 - -

E

5

4 -

3 -

2 -

1 ■

r£

E k 150

D3 1/0.SE*03

nDz0.3A2n3o

_DL N/mm'

-4-50 100 A S

20 0 250

Figure 5.6 : De ta i l t e s t e d at T.R.R.L.

92 -



Inf luence line

0. 5 -■

Histogram Rain f low

NI /NTOT

Rheden Traf f i c

0. 4 -■

0. 3 - •

0.2 - •

0. 1 - •

100 20 0 300 ¿00

kN

500

Fatigue damge M D I

1 0 - •

g . .

8 - -

7 ■ -

6 - -

S ■ -

¿ - -

3 ■ -

2 -■

1 ■-

=d£ i L

D3 1 / 0 . 1E .0 5

n m =0.108 na j ,

UL . kN - H ■ -SOO 00 20 0 3 00 ¿0 0

Figure 5.7 : Detai l t es ted at LBF

93



Hydraulic actuator

Reaction frame

Figu e 5.8 : Fatigue t e s t rig

94



Soirth si d«

TYPE A SPECIMEN

South si dì

Fatigue crack at toe of velo

Figure 5.9 : Location o f cr ack s in fat ig ue t e s t specimens

95



lOÜO

STRESS RANGE N/mm 2

(O O)

100

Mean line 9Ş% confidence Limirs

o crack a • crack b x crack c

Eurocode class 125

Eurocode class 50

J l i l i

CYCLES I U '

Figure 5.10 : Fatigue t e s t s a t cons tan t ampli tude - TYPE A* SPECIMtNS



CO

1000'

3 0 0 -

HANGE

( l l / mm a )

1 0 0

50

10-

APPLJED «IMÍCTUWM

"1 1 1 1" I i l l

Ş-ÎLii'Unviîş rem

Evnocohiz cLÃãs 'uu

i 1—i—i—i i i i

[-lean l i n e

V57. c o n f i d e n c e 1 imi LK

Test::» a t c o n f i a n t . ;>mi> I. ¡ L u d e T e s i s a t v a r i a b l e amp i I U n i e

10

C YC LES

Figure 5.11 : Tests on TYPE B SPECIMEN



TYPE S

Crack

TYPE T

Erack

TYPE R

Crack

Figure 5.12 : Location of the cracks

in the fat igue tests specimens

- 98



TROUGH TO CROSSBEAM CONNECTION OF ORTHOTROPIC STEEL BRIDGE DECK

CO CO

i o c -7

ro

° -H m «^f ro -•-

OJ » O

<D <">

■ a n> ro o

S'Y1

*~ I

ro a f ro<" .— i - h -J

-*■ CU 3 UD CD

A

I

I

n.

< o j CD

(/) V) O f t U

2 O K

m

Lu

O

z: <

co co IJ J

ce i—

co

400

360

320

280

240

200

180

160

140

120

100 -

90

80

70

60

50

(S-N Curves according lo EUROCODE 3)

1— r I I I I I I I T "I—I I I I I II

40

i i— i i 11 i i 1—i—n-rrru CONSTANT AMPLITUDE TESTS E

THEORETICAL NOMINAL STRESS RANGE -

TUD-ECSC Resea r ch

•. TYPE - T : Con n e c t i o n w i t h ' o l d * c u t - o u t «

A TYPE -R: Con n e c t i o n w i t h 'new' c u t - o u t «

C O í ^ 5 C O ô T Y P E - S : Con n e c t i o n w i t h o u t c u t - o u t «

TU -DELFT S t e v i n LABORATORY S t e e l STRUCTURES

I l I l l LLU i.

— r u n - o u t

3

10 3 4 5 6 7 9

J I I I I I II .J I I I I I I I J l_.L

3 4 5 67 9 _U±

10 10 3 4 5 6 7 9 7

10 3 4 5 67 9

10

NUMBER OF CYCLES (AT VISUAL CRACK INITIATION) [N ] CRB.GRA



o o

UD

c ro u i ~ & ■p-

- f o

s« QJ UD

UD -* •

ro < " - * ■

50

ro

A

o.

UJ

o

iÜ

3 T ro

- s w -J c o «—

s ir

,?-> TD

£ M

a" 1. O ^ ro Q -

o K. u.

> ■ij

fM

*m^

LÜ

<J )

-z .

er C O C O Ld

o : r -co

TROUGH TO CROSSBEAM CONNECTION OF ORTHOTROPIC STEEL BRIDGE DECKS

400

360

320

28 0

24 0

20 0

18 0

16 0

140

12 0

10 0

90

80

70

60

50

40

( S - N Cu r v e s a c c o r d i n g t o EUROCODE 3 )

i n - n r m 1—i—r-rrnr

= wel

//////////s.

12

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JU-DELFT S t e v i n LABORATORY S t e e l STRUCTURES

I l l I I LLLL

"i i r~i i 11 n i — t — i - i n r n T CONSTANT.AMPLITUDE TESTS E

LOAD RANGE 2 00 - 26 8 KN

TUD-ECSC Resea r ch

U TYPE -T : Con n e c t i o n w l l h ' o l d ' e u t - o u t i

A TYPE -R: Con n e c t i o n w i t h ' n ow ' cu t—out»

-•»run—oul

3 4 5 6 7 9 J I I

10 • 1 0

NUMBER OF CYCLES

3 4 5 6 7 9 J I I I I I I I I 2 3 4 5 6 7 9

J _ J _ _ L . L J JX l

1 0 10

(A T A CRACK LCNGTH OF 50 MM)

2 3 4 5 6 7 9 o 10

[N ] >

CRU-12 .GRA



en

u u

QLU

e noÜ_

e nLUCL

LO

LO

dl

DI



load

oto

crack no. 1 after 205 000 cyclescrack no. 2 after 339 000 cyclescrack no. 3 after 339 000 cycles (no crack growth observed )crack no. U after 3 U 000 cyclescrack no. 5 after 350 000 cyclescrack no. 6 after 350 000 cyclescrack no. 7 after 350 000 cycles

T : region with tension stressesC : region with compression stresses

Figure 5.16 : RESULTS OF CONSTANT LOAD AMPLITUDE TEST WITH A SPECIMENHAVING CUT-OUTS OF THE SHAPE PROPOSED TO BE BETTER (TYPE I )



load

oco

crack no.crack no.crack no.

crack no.e n d of test

afterafterafter

afterafter

2 8U 2003 ¿09 900¿ 5¿1 000

5 780 3005 926 600

cyclescyclescycles

cyclescycles

T : reg ion wi th tension s t ress esC : reg ion wi th compression st resses

Figure 5.17 RESULTS OF THE VARIABLE LOAD AMPLITUDE TEST WITH THE SPECIMENHAVING CUT-OUTS OF THE NOW USUALLY USED SHAPE (TY PE I I )



load

o

crac k no. 1 after 6 728 000 cyclescrack no. 2 a f te r 6 728 000 cyc les (no crack growth obse rved)crack no. 3 a f te r 6 728 000 cyc les (no crack growth obs erved)crack no. U af ter 6 728 000 cyc les (no crack growth obser ved)crack no. 5 a f te r 13 385 000 cyc les (no crack growth ob serve d)crac k no. 6 after 16 089 000 cycle send of test after 22 129 000 cyc les

T : reg ion wi th tension s t resse sC : reg ion wi th compression st resses

Figure 5.18 RESULTS OF THE VARIABLE LOAD AMPLITUDE TEST WITH A SPECIMENHAVING CUT-OUTS OF THE SHAPE PROPOSED TO BE BETTER (TYP E I )



oUI

8 0 0

kN

¿ 8 5

¿ 0 0

-ooo

enca

200

100

I I I I I l i l i

S-N t es t resu l t sP = 500 kN

15 kNu=

U c r a c k s

equ iva len t l oad

E n , A P MA P K =

Un, A Pi"

e s t i m a t e d S - N c u r v eslop e k = 5

I I i I U l l 1 — T

resu l t s o f va r iab lea m p l i t u d e t e s t s

Pmax = 500 kNP.. = 15 kN

l i l i l í

B-¿-¿ - E ¿

e-

k = 5k = A

-E¿-k = 3

equ iva len t l oad cyc les

E n , A P :k

i *- ' i

E A R '

1— I I I H i l l 1 — ' ' ' m i i i i i i M M I i i i i i n n10 4 2 3 4 56 8 1 0 5 2 3 4 56 8 1Q 6 2 3 4 56 8 1Q 7 2 3 4 56 8 i n 8

10 (

c y c l e s

Figure 5.19 : RESULTS OF THE CONSTANT AND VAR IABLE LOAD AMPLITUDE TE STS (TYPE I)



6 . O R T H O T R O P I C D E CK T O CR O S S B E A M C O N N E CTI O N .

6.1. Introduction.

The bolted connections tested are used to attach orthotropic deck

plates to supporting I section crossbeams. This type of construction isfound in harbow arens where the bridges carry both road and rail traffic.

These orthotropic steel decks have to be dismountable.

These bolted connections are subjected to complex forces resulting

from angular rotation of the supports (lever effect), from vertical forces

(tension or compression) and from horizontal strains (axial shortening).

These forces are difficult to calculate because they depend on element

stiffness and secondary effects.

Tests were carried out on different specimen configurations

subjected to forces representing those found in bridges.

In addition, tension-bending tests on bolts were carried out.

6.2. Results from the University of Liège on the connection.

The problem was solved using two approaches :

* introduction of fexible elements to reduce secondary bending

stresses.

Tests gave the solution presented in Fig. 6.1. ;

* using elements with sufficient stiffness to reduce secondary

effects.

6.3. Results of fatigue tests on 8.8 bolts.

These tests were carried to study the influence of the following

parameters on the fatigue strength of the bolts :

* stiffness of the beams with influences the angular rotation of bolt

heads and thus influences the bending stresses in the bolts ;

* prestressing ;

* introduction of neoprene element ;

Characteristics of tested bolts are :

* nominal diameter : 12 mm ;

* measured net section : 8*4,3 ram2

* measured ultimate stress : 896 N/mm 2

* Brunell hardness : 280 to 295.

- 106 -



Test beams are :

* series 1 : IPE 160 - without prestressing.

* series 2 : IPE 160 - with prestressing.

* series 3 : IPE 160 - with neoprene and steel element.

* series H : HEM 120 - without prestressing.

* series 5 : HEM 120 - with prestressing.

A specimen of series k is shown in Fig. 6.2. and test results are

presented in Fig. 6.3»

Prestressing has favourable effects of reducing the angular

rotation and the stress variation.

Figure 6.3. shows that the Eurocode 3 classification is

conservative. It is lower than the most unfavourable tests (flexible

beams without prestressing).

Before to have general conclusion, more test are necessary.

- 107



o00

-Ó

]max :80kN

Q min : 5 kN

1363

Figure 6.1

Coupe. A A

pla t20 5

12

-m a smaamaa t

raidisseur

L_ 100/100/10

rondel leneoprene frette 6Q/3Q/7

ctìT\ neoprene frette 80/80/7



t

70mm

rondel le d'acier 024

i »

♦

rondel le d'acier#24 ep 2mm

point de rupture du boulon

Figure 6.2

109



Boulon 8.8-012

400

32 0

— 2 8 0Cu

CL 2 4 0

CO

2 00

Ü JCD

ŒÛC

COCOU J

enh -m

i 60

mo

12 0

10 0

80

70

60

5 0 -

40

10

I i I I I 1 1 1 I T T

Í.DX

5A

i i i M m I 1 I I I I I I

T T

X5B

SC

AC• • •—*-

3A 3D

3C

t B

2A

3 4 5 6 8 3 4 5 6 8

Figure 6.3

i i i i i i n I i i i i i m3 4 5 6 8 2 3 4 5 6 8

1 pr ofi lé IPE 160

3 prof i lés IPE 160»néoprène

4 prof i lés HEM 120

2 profi lés IPE ICO -p r e s t r e s s e d b o l d .5 pr o fi ls HEM 120 -

N.B. : for s er ie s 2 and 5d rawn s t r e s s r an g esa re n o t a c tu a l s b u tare obta ined as ifbolts where notp r e s s t r e s s e d .

10CYCLES N

10



7 . APPLICATIONS.

The sponsors of this research are the European Community for Cool

and Steel, the National Autorities responsible for bridge building and

bridge builders.

All the results obtained in the laboratories participating to the research

have been widely disseminated in each country . Some of the information is

already used in the new versions of the design rules and national

recommandât ions.

The fatigue strength of the connection stiffener-plate given in

the Belgian Standard NBN 5 has been changed, taking in account the new

results. The experts busy with the redaction of Eurocode 3, have also

been informed of the new results.

The conclusions of chapter 3, concerning the minimum value of the

root gap of the connection stiffener-stiffener is now a rule for the bridge

builders in Netherlands, Italy and France. Designer welder and superviser

are informed, and pay particular attention to this very important detail.

In Germany, the new shape of the cut-outs in the web of the

cross-beam is a part of the design rules of Deutsche Bundesbahn following

the investigations of Haibach and Plasil [2 6]. The research gives

informations concerning the fatigue behaviour of this shape in roadbridges.

Orthotropic decks are excluded form the UK bridges design code

[11] but the code advises the designer to seek specialist advice. That

advice is frequently provided by the TRRL, so that the results of research

are passed on directly to bridge designers. In addition, designs are

subject to technical approval by the Department of Transport who receive

report of all TRRL research.

The results have already been applied in some bridges in

construction.

The new automatic welding procedures studied by IRSID and LCPC for

the connection stiffener-plate have been applied to the fabrication of the

orthotropic steel deck of two large bridges in France :

Pont de Cheviré on the Loire river :

Pont de Normandie on the Seine rive.

Submerged arc welding was used, without edge preparation and the

lack of penetration below 1 mm.

- 111 -



Analytical methods developed enable the stress range in the

stiffener-plate, weld to be calculated for various stiffener shapes.

Following this calculation, the shape of the stiffener was changed in the

design of the Kruiskansburg in Antwerpen (reducing of the thickness of the

stiffenerweb and increasing of the height) in such a way as to increasesignificatly the fatigue life.

The results obtained at the University of Pisa for the

stiffener-stiffener connection have been used in the preliminary studies

for the design of the Messina Bridge in Italy (suspended steel bridge, with

a span of 3300 metres and with six lanes for road traffic and two rail

ways).

The results obtained by the testing of the bolted connection ofthe orthotropic deck to cross-beam are used for the design of repairing the

Oosterweelburg in Antwerpen and the design of the new Kruiskansbridge.

8. CONCLUSIONS.

The first two phases of this common reserach programme were

concerned with the collection of traffic data needed to determine loads on

road bridges. This third phase concerned the fatigue behaviour oforthotropic steel decks.

It was necessary to analyse the fatigue behaviour of all the

connections in the orthotropic deck. Because of the scale of the problem,

it was not possible for all the testing and analysis to be carried out in a

reasonnable timescale in one laboratory or one country. A common research

programme was therefore formulated and the work distributed between seven

laboratories situated in six countries of the European Community.

For some connections improvements in the design are suggested :

for others, unexpected behaviour occurred which require further

investigations.

Stiffenei—pkate connection.

A welding procedure has been defined which gives a much better

fatigue resistance and excludes cracks in the weld.

- 112



The calculation method developed allows alternative stiffener

shapes to be assessed which may lead to an improved deck design.

A fracture mechanics model has been developed and calibrated using

the experimental results. Such a model, which can be used on a microcomputer, has already been used successfully to study fatigue problems on

existing bridges.

Stiffener-stiffener connection

The tests on the stiffener-stiffener connection with backing

strips showed the importance od the root gap before welding and indicated a

better behaviour of triangular compared with trapezoidal stiffeners.

Triangular stiffeners butt welded from both sides dit not give the

expected improvement in fatigue strength over the butt weld made from one

side on a backing strip.

This important information can be directly applied by bridge

builders. However, further investigations are required to clarify these

results and assess the influence of construction procedures and residual

stresses.

Fatigue cracks are expected to develop in this connection on

existing bridges in a short time. Repair procedures should therefore be

developed as a matter of urgency.

Stiffener crossbeam connection.

Test results from the stiffener-crossbeam connection indicate a

minimum fatigue classification of Eurocode class 80. However, cracks

occured where they were not expected and more information is required on

the shape of the cut-outs in the crossbeam.

A cut-out around the apex of the triangular rib gave an

improvement in fatigue strength over a welded all round connection.

However, cracks developed unexpectedly at small cut-outs near the deck

plate.

Conversely, for trapezoidal ribs, cout-outs around the apex of the rib

reduced the fatigue strength, the "optimised" shape of cut-out proposed for

Germam railway bridges dit not give the expected improvement in fatigue

strength in these tests.

113



Before new tests are carried out, a computer analysis of the

stress distribution at the critical points is necessary to explain this

behaviour an to assess new shapes of this connection.

Bolted connections.The fatigue tests on bolted connections show that the

classification given in Eurocode 3 corresponds to the worst case.

This connection, which occurs frequently in steel bridges, needs further

investigation in order to define more realistic S-N curves.

Results from tests carried out under variable amplitude loading

generally showed good agreement with lives estimated using the Miner rule.

It may be concluded that in furture, fatigue tests can be carried

out mainly under constant amplitude loading. These tests are more

economical. Only a small number of tests under variable amplitude loading

are then required to confirm the application of the Miner rule for each

particular case.

This work gives a lot of information on orthotropic steel decks

which can be used directly by bridge designers and builders. However, some

problems remain which require furhter investigation.

114



BIBLIOGRAPHY

[1] Dr. Lehrke

"Messung und Interpretation von Dynamischen Lasten an StahlbrUcken".

Forschungsvereinbarung Nr. 7210-KD/119 - Abschlussbericht - Phase 3.

Fraunhofer - Institut rur Betrieksfestigkeit.

[2] A. BRULS - E. POLEUR.

"Mesures et Interprétation des Charges dynamiques dans les Ponts".

Recherche n° 7210-KD/201 - Rapport final - Phase 3.

Service Ponts et Charpentes, Université de Liège.

[3] A. BIGNONNET (IRSID) and CARRACILLI, B. JACOB (L.C.P.C.)

"Mesures et Interprétation des charges dynamiques dans les Ponts".

Recherche n° 7210 - KD/317 - Rapport final - Phase 3.Institut de Recherches de la Sidérurgie Française.Laboratoire Central des Ponts et Chaussées.

[H] S. CARAMELLI, P. CROCE, M. FROLI, L . SANPAOLESI.

"Misure ed Interpretazioni dei Carichi dinamici sui Ponti".Convenzione n°7210 - KD/¿)11 - Relatione technica finale - Fase 3

Instuto di Scienza delle Costruzioni, Università di Pisa.

[5] H. KOLSTEIN, J. de Back.

"Mesurements and Interpretation of dynamic loads in Bridges".Agreement Number : 7210-KD/609 - Final Report - Phase 3.

Delft University of Technology - Stevin Laboratory.

[6] C. BEALES.

"Measurements and Interpreation of dynamic Loads in Bridges".

Agreement Number : 7210 - KD/807 - Final Report - Phase 3.

Transport and Road Research Laboratory.

[7] DE BACK, A. BRULS, J. CARRACILLI, E. HOFFMANN, L. SANPAOLESI, J.P. TILLYand J.M. Zaschel.

"Measurements and Interpretation of dynamic Loads on Bridges".

Synthesis Report - Phase 1.

Commission of the European Communities.

EUR 775^ FR, EN, DE (1982).

[8] E. HAIBACH ,J. De BACK, A. BRULS, J. CARRACILLI, B. JACOB, M.H. KOLSTEIN,

J.PAGE, M.R. PFEIPER, SANPAOLESI, TILLY, ZACHEL, HOFFMAN.

Measurements and Interpretation of dynamic Loads on Bridges".

Synthesis Report - Phase 2.Commission of the European Communities.

EUR 9759 FR, EN, DE (1986).

- 115 -



[9] A. BRULS, B. JACOB, G. SEDLACEK and all.Traffic data of the European countries.Eurocode on action - Part 12 : Traffic loads on Bridges.Report of working groupe 2.

[10] NBN5 "Ponts en acier - projet 1987".

Institut Belge de normalisation, Bruxelles.[11] BS 5400 - Steel, concrete and composite bridges. Part 10 : Code of

practice for fatigue. British Standards Institution, London, 1980.

[12] Eurocode n°3 : Design Steel Structures. Final Draft. December 1988.

[13] NEN 2063. "Arc Welding - Fatigue loaded structures - Calculation ofwelded joints in unelloyed and low-alloy steel up to an including Fe510 (Fe 52)".Nederlands Normalisatie Instituut march 1988. Delft.

[14] MADD0X S.J.

"The fatigue behaviour of tropezoïdal stiffener to deck plate welds inorthotropic bridge decks".T.R.R.L. Supplementary Report 96 U.C.

[15] TH0NNARD - JANSS."Comportement en fatigue des dalles orthotropes avec raidisseurstrapézoïdaux.CRIF : Section Métallique - MT lol - Août 1985.

[16] D.E. NUNN AND J.R. CUNINGHAME.

"Stresses under wheel loading in a steel orthotropic deck with VStiffeners".

TRRL Report 59 U.C.

[17] D.E. NUNN AND J.R. CUNINGHAME.

"Stresses under wheel loading in steel orthotropic deck withtrapézoïdal stiffeners".TRRL Report 53 U.C.

[18] HUber M.H. "Die Grundlagen einer rationellen Berechnung der kreuzweisebewerthn Eisenbetonplatten". Zeitschrift der Osterreiches Ingenieurund Architekten-Vereines, n.30, 1914.

[19] Pelikan W., Esslinger M. : Die Stahlfahrbahn, Berechnung undKonstruction". M.A.N. Forschungsheft, 7, 1957.

[20] CROCE P. : "Linfluenza della variabilità di sezione nelle nervature diuna piastra ortotropa". Atti dell'Istituto di Scienza delleCostruzioni, n.252, vol. XVIII, Pisa, 1988.

[21] LEVY M. : "Sur l'équilibre élastique d'une plaque rectangulaire".Compte rendu acad. S c , 129 1989.

[22] CUNINGHAME J.R.Steel bridge decks : Fatigue performance of joints betweenlongitudinal stiffeners. TRRL. Report. LR 1066.

116



[233 K. YAMADA

Fatigue Strength of Field-Welded Rib Joints of Orthotropic SteelDecks.IIW. Doc. XIII 1282-88.Department of Civil Engineering, Nagoya University.

[24] TROMP : Fatigue of field splices in ribs of orthotropic steel bridgesdecks. Report 6-7*1-15.Stevin Laboratory - Steels structures - Delft University ofTechnology 197*1.

[25] A. BRULS

Mesures et interprétation des charges dynamiques dans les ponts.2ème phase. - REcherche CECA. Rapport EUR. 8864.

[26] E. HAIBACH, PLASIL.

Untersuchen sur Betriebsfestigkeit von Stahlleichtfahrbahnen mit

Trapezhohlsteifen im Eisembahnbrückenbau.Der Stahlbau 53~S 269-27*1. 1983.

117



European Comm unities — Commission

EUR 13378 - Meas ureme nt and interpretation of dynamic loads inbridgesPhase 3: Fatigue behaviour of orthotropic steel decks

A. Bruls

Luxem bourg: Office for Official Publications of the European Com munities

1991 - VI, 117 pp., num . tab., fig. - 21.0 x 29.7 cm

Technical steel research series

ISBN 92-826-0532-9

Catalogue number: CD-NA-13378-EN-C

Price (excluding VAT) in Luxembourg: ECU 10

This research, carried out with the financial help of the European Coal andSteel Community, concerned the fatigue strength of orthotropic steel decksof road bridges. It followed two phases that were con cerned with the collection of traffic data and the me asurement of stresses produce d in bridges. Fatigue tests under constant and variable amplitude were carried out on stif-

fener-plate co nnections, stiffener-stiffener connec tions with U and V shapes,and stiffener cross-beam connec tions. From the test results and calculationssome co nclusions can be drawn which are directly usable in bridge design.However, some unexpected behaviour occurred and some connectionsneed further investigation.


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Fatigue Behaviour of Orthotropic Steel Decks

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