Steel Construction 01-2012 (sample issue)

Steel ConstructionDesign and Research

– Hybrid steel plate girders subjected to patch loading– New resistance model for plate girders under patch loading– Design of trapezoidal corrugated web girders under combined actions– Intermediate transverse stiffeners in plate girders– Reduced stress design of plates under biaxial compression– Innovation brokers for managing interdisciplinary research– China’s longest single-pylon cable-stayed bridge span– Reconstruction of “Olympic” stadium in Kiev

1Volume 5Februar 2012ISSN 1867-0520

01_SC_U1_1-12_EinzelnNEU_Layout 1 20.02.12 13:51 Seite 1

Form + Function = DETANDETAN tension rod systems. Your solution for transparent design.

Many advantages with one result: HALFEN-DEHA provides safety, reliability and effi ciency for you and your customers.

F orm and function are perfectly combined in the DETAN tension

rod system. Individual system solutions make it possible to realise even the most complicated designs and aesthetic details both indoors and outdoors.

DETAN tension rod structures can consist of one or all of the standard components shown in this example.

SimpleWith screw connections, no welding is required. Standard tools mean simple and convenient installation.

SafeDETAN stands for reliable planning through certifi cation, type testing and dimensioning software. A high load capacity allows safe use in a wide range of applications and gives more design freedom for aesthetic architecture.

Effi cientMaterials such as carbon steel S 460 and cast steel for the fork ends have signifi cantly increased the effi ciency

of the tension rod system. The high strength of the tension rods makes them thin and saves on material use: cost optimisation and design in a single stroke.

anchor disk

locking nut

seals (optional)

coupler with sailtension rod

connection plate to structure

locking nut

fork endpincirclip

HALFEN GmbH · Liebigstr. 14 · D-40764 Langenfeld · Tel.: +49 (0) 2173/970-0 · www.halfen.comHALFEN GmbH · Engineering Support · Tel.: 02173/970-90 35 · www.halfen.de

02_SC_U2_1-12_02_SC_U2_2-09.qxd 09.02.12 12:06 Seite 1

Content

Steel Construction1

Editorial

11 Ulrike KuhlmannEurocode 3 Part 1.5: Plated Structures – new chances and developments

13 Rolando Chacón, Marina Bock, Enrique Mirambell, Esther RealHybrid steel plate girders subjected to patch loading

10 Rolando Chacón, Benjamin Braun, Ulrike Kuhlmann, Enrique MirambellStatistical evaluation of the new resistance model for steel plate girders subjected topatch loading

16 László Dunai, Balázs Kövesdi, Ulrike Kuhlmann, Benjamin BraunDesign of girders with trapezoidal corrugated webs under the interaction of patchloading, shear and bending

23 Franc Sinur, Darko BegIntermediate transverse stiffeners in plate girders

33 Benjamin Braun, Ulrike KuhlmannReduced stress design of plates under biaxial compression

41 Gregor Nuesse, Mukesh Limbachiya, Roland Herr, Alan EllisManaging interdisciplinary applied research on sustainability in construction withthe help of an innovation broker

Reports

53 Yingliang Wang, Chengzao Huang, Yuncheng FengHuangpu Pearl River Northern Channel Bridge, Guangzhou – the longest single-pylon cable-stayed bridge span in China

61 Aleksandr V. ShimanovskyAwaiting the 2012 European Football Championship: some features of the reconstruction of the stadium of the “Olympic” National Sports Centre in Kiev

Regular Features

19 Peoples15 News (see 40, 60)52 Conferences66 ECCS news

A4 Products & Projects

The Huangpu Pearl River Bridge is the key project on the second ring highway ofGuang zhou, China. The main bridge over the northern channel is a single-pyloncable-stayed bridge with a main span of 383 m and a steel box girder, which wasconstructed using the balanced cantilever method (see p. 53).

Volume 5February 2012, No. 1ISSN 1867-0520 (print)ISSN 1867-0539 (online)

Wilhelm Ernst & SohnVerlag für Architektur und technische Wissenschaften GmbH & Co. KGwww.ernst-und-sohn.de

Wiley InterScience is now Wiley Online Library

www.wileyonlinelibrary.com, the portal forSteel Construction online subscriptions

Journal for ECCS members

03_SC_Inhalt_1-12_03_SC_Inhalt_1-09.qxd 10.02.12 08:02 Seite 1

A4 Steel Construction 5 (2012), No. 1

Products & Projects

ing damper mass of 150 kg each is separated into several steelplates such that it can be adjusted retroactively. The plate stackrests on four steel coils which are mounted on a base plate. Thebase plate is bolted to the longitudinal girders. For each damper,the frequency of the coils is tuned to the natural frequency ofthe platform and lie between 2.03 and 2.68 Hz.

Coils plus hydraulicsHowever, the countereffect of the swinging mass plates alonewould not suffice to damp the structure against vibrations whichare created by groups acting willfully. For this reason, Maurerintegrated in the centre of the mass damper an additional hy-draulic damping element which is connected with the mass andacts as internal brake. It converts motional energy into heat, andthis way structural vibrations decay very fast. In comparison tothe “undamped” state (i.e. without hydraulic damping element)it not only takes more people to trigger a perceptible vibration,but the vibration also decays in the course of a few seconds. The four mass dampers were installed from above in summer2011, by way of detaching the grid irons, lifting the damper boxesto their location and fixing them. After the inauguration, finalmeasurements were made as well as a fine adjustment of thedamper mass.

Further Information:Maurer Söhne GmbH & Co. KG, Frankfurter Ring 193, 80807 München, Tel. + 49 (0)89 32394-0, Fax + 49 (0)89 32394-234, [email protected], www.maurer-soehne.de

Tamed, but not tameThe name AlpspiX depicts a spectacular viewing platform about1,000 m over the Hell Valley, being opposite of Germany’shighest mountain, the Zugspitze. Two light platforms protrudein the shape of an X into the abyss. However, in case of pur-poseful movements of the visitors, these platforms vibratedthat much that the operator, the Bayerische Zugspitzbahn Berg-bahn AG, decided to make the platform also accessible totourists which are more sensitive to vibrations. The “Dampingwith residual adventure factor” was implemented by MaurerSöhne, the world wide renowned specialist for structural pro-tection systems.

The inauguration of the AlpspiX platforms was in July 2010. Theyare located at the “Osterfelderkopf”, which is only 50 m abovethe mountain terminus of the Alpspitz Track, and thus they alsocan be reached by inexperienced hikers. This in contrast to theAlpspitz peak which is 550 m further up and which will usuallybe reached by means of a fixed rope route.The AlpspiX viewing platform facilitates an undisturbed view atits finest. Only a glass plate prevents the free fall at the tip of theplatform. And a grid iron allows a look into the abyss 1,000 mbelow. But that the steel platform were prone to undamped vi-brations was not exactly palatable for many visitors. Already afew individual visitors created continuing vibrations in the plat-forms – not to mention the jaunty groups who purposely createdthese vibrations.

Dampen, but not fixateThe job instruction for the damper experts of Maurer Söhne wastherefore to reduce the vibrations but not completely eliminate,by way of tuned mass dampers. Next to the higher attractivenessof the visitors the owner also considered the service life of theplatforms: the less vibrations, the less material fatigue, and thuslonger maintenance intervals and longer service life.The two AlpspiX arms which are each about 17.50 m long aresituated on top of each other and protrude in the shape of an Xinto the abyss. Their load carrying structure consists of two eachdouble-T-girders with cross members. The two platforms are sep-arately constructed and clamped into two foundations each.The tuned mass dampers had to be installed at the tip of thecantilever, because there the vibrations are the strongest. Further-more, the 2 tuned mass dampers each should not disturb the freeview. Therefore, the 340 mm high dampers were placed underthe grid iron, between the longitudinal girders at the tip of theplatform.The four tuned mass dampers of each 340 kg have the shape offlat square boxes with a footprint of 650 × 650 mm. The swing-

A W i l e y C o m p a n y

Bild 1. AlpspiX: spektakuläre Aussichtsplattform 1.000 m über dem Höllental und ge-genüber der Zugspitze

Bild 2. Querschnitt durch Massendämpfer: Die viereckigen, flachen Boxen (orange) be-herbergen die schwingende Tilgermasse (grün) von je 150 kg, die in mehrere Plattenaufgeteilt ist. In der Mitte ist ein zusätzliches hydraulisches Dämpfelement (gelb) inte-griert. Der Plattenstapel liegt auf den vier Federn (blau) auf, welche auf der Boden-platte (rot) befestigt sind. Die Befestigung am Bauwerk erfolgt mittels vier Schraubenzu den Doppel-T-Trägern. (Foto/Grafik: Maurer Söhne)

04_SC_A04-A10_1-12_SC 09.02.12 12:13 Seite 4

It was in summer 2011 that Ceram KoteInternational GmbH launched the inno-vative, solvent-free coating product PRO-GUARD CN-OC with its excellent pro-tective properties. Extensive test seriesin cooperation with an independent re-search institute as well as relevant prac-tical experience verify the outstandingperformance of PROGUARD CN-OC.

The new solvent-free coating product hasan excellent chemical resistance. Thatmeans stability against: – various acids and alkaline solutions

(pH value 4–11), in some cases up tooperating temperatures of 110 °C,

– all kinds of hydrocarbons, and– “killer-solution” – already > 8000 h

(> 11 months!) of uninterrupted storagein 98 % sulphuric acid, pure methanoland 3 % sodium chloride solution witha mixing ratio of 1:3 [1]in each case anda constant temperature of 50 °C.

Furthermore, it has phenomenal adhe-sion:– on concrete (direct application without

the need for a primer first is possible),– on carbon steel, and – on stainless steel.Remarkable results are achieved whencuring at room temperature and there isa great increase in adhesion at elevatedtemperatures. The test procedures and re-sults are described below.

Adhesion tests withPROGUARD CN-OC To identify the adhesion value, shearstrengths were tested with specimens ofstainless steel (see Fig. 1). Firstly, cylindersand base plates were sandblasted, then acoating of PROGUARD CN-OC was ap-plied in the centre of the base plates andthe cylinders placed on these without ap-plying pressure. The coating was allowedto cure at room temperature for 24 h. Af-terwards, different tempering steps werecarried out to examine the behaviour ofthe coating under thermal influences. The prepared specimens were sheared offwith an “Instron 4469” tensile testing ma-chine. The determination of bondingstrengths included different analyses of

shear strengths (average of five specimensof each kind) and the fracture behaviourof the coating.

Interpretation of test results The progression of adhesive strength onstainless steel for PROGUARD CN-OCis highly interesting from the scientificviewpoint. After curing at room tempera-ture, the adhesion on stainless steelreached an average value of 17.7 MPa.Tempering at 100 °C increases the adhe-sion dramatically by 74 %.A cohesion fracture occurred in the coat-ing at the average value of 30.7 MPa.That means the internal strength of PRO-GUARD CN-OC in shear tests is definedby this value. The true adhesion of thecoating on stainless steel can be consid-erably higher. To assess the actual bond-ing strength, further tensile tests with otherprocedures would be necessary. Whatcauses this extreme increase in adhesionon stainless steel? More surface analyseswere carried out to obtain the scientificexplanation for this phenomenon.

Surface analyses of stainless steelsamples using x-ray photoelectronspectroscopy (XPS)Preparation: All the base plates of thestainless steel specimens from the previousshear tests were cleaned in N-methyl-2-pyrrolidone (NMP) at 60 °C and for 72 husing an ultrasonic process to remove thecoating thoroughly. The plates were coatedbeforehand only in the area of the affixedcylinder. Thus, a direct comparison be-tween the coated and uncoated metalwas possible to define the effect of PRO-GUARD CN-OC.Test apparatus: “Theta Probe Thermo VGScientific”, 400 mm region of examination,10 nm penetration depthAssay method: (source: Wikipedia[2])X-ray photoelectron spectroscopy (XPS)is a method by which the chemical com-position of solid objects and their surfacescan be determined. It uses qualitative ele-mental analysis, i.e. it reveals the chemicalelements contained in a solid object. Merelyhelium and hydrogen cannot usually be de-tected because of their low cross-sections.

Products & Projects

Application of organo-ceramic coating product gives phenomenal adhesion on stainless steel

Fig. 1. Shear test in accor-dance with DIN EN 1465

04_SC_A04-A10_1-12_SC 09.02.12 12:13 Seite 5

XPS analysis findings – result of examination, exemplary after tempering at 100 °C for 4 hThe different C1s spectra of the two surfaces demonstrate un-equal chemical carbon constitutions in terms of their oxidationstate or molecular character.


Products & Projects

*Pr

ices

are

val

id in

Ger

man

y, e

xclu

sive

ly, a

nd s

ubje

ct t

o al

tera

tions

. Pric

es in

cl. V

AT.

Boo

ks e

xcl.

ship

ping

. Jou

rnal

s in

cl. s

hipp

ing.

020

0100

006_

dp

Ernst & SohnVerlag für Architektur und technischeWissenschaften GmbH & Co. KG

Customer Service: Wiley-VCHBoschstraße 12D-69469 Weinheim

Tel. +49 (0)6201 606-400Fax +49 (0)6201 [email protected]


Steel Structures The Finite Element Method (FEM) has become a standard tool used in everyday

work by structural engineers having to analyse virtually any type of structure. After a short introduction into the methodolgy, the book concentrates on the

calculation of internal forces, deformations, ideal buckling loads and vibration modes of steel structures. Beyond linear structural analysis, the authors focus on various important stability cases such as fl exural buckling, lateral torsional buckling and plate buckling along with determining ideal buckling loads and second-order theory analysis. Also, investigating cross-sections using FEM will become more and

more important in the future. For practicing engineers and students in engineer-

ing alike all necessary calculations for the design of structures are presented clearly.

Author information:Univ.-Prof. Dr.-Ing. Rolf Kindmann teaches steel

and composite design at the Ruhr University in Bochum and is a partner of the Ingenieursozietät Schürmann-Kindmann und Partner in Dortmund.

Dr.-Ing. Matthias Kraus is a research assistant at the same chair.

R O L F K I N D M A N N ,

M AT T H I A S K R A U S

Steel StructuresDesign using FEM

2011. 54 pages.

. Softcover. 59,– *

ISBN 978-3-433-02978-7

Ord

er

on

lin

e:

ww

w.e

rnst

-un

d-s

oh

n.d

eStainless steel, uncoated (red graph):– The carboniferous film on the surfaces caused by environmen-

tal influences features a constant chemical carbon constitutionirrespective of the tempering.

Stainless steel, coated beforehand (green graph):– The organo-ceramic components of the coating adhere to the

surface and cannot be removed by ultrasonic cleaning. – The carboniferous film on the surface also exhibits other

chemical carbon constitutions in comparison with environ-mental carbon sediments.

– By increasing the tempering, one kind of chemical carbon con-stitution increases and another kind decreases proportionally.

Conclusion The scientific surface analysis shows that the uniquely formulatedhigh-performance coating PROGUARD CN-OC contains partic-ular, innovative organo-ceramic fillers that form specific organiccompounds even with slight tempering. These compounds causechemical interactions with compounds in the surface of themetal and therefore give rise to extreme adhesion on stainlesssteel substrates.Currently, this chemical behaviour of PROGUARD CN-OC isbeing explored in depth. We are also carrying out research todetermine at which tempering thresholds the chemical reactionsof both materials occur in order to identify the highest possibleadhesion of PROGUARD CN-OC on stainless steel.

Further information:Ceram Kote International GmbH, Daimlerring 9, 32289 Rödinghausen, tel. +49 (0) 5223 96276-0, fax +49 (0)5223 96276-17, [email protected], www.ceram-kote.de

Fig. 2. Interpretation of test results (© Ceram Kote International)

Adhesive shear strengths in MPa

04_SC_A04-A10_1-12_SC 09.02.12 12:13 Seite 6

Aluminium design to Eurocode 9 with RSTAB and RFEM

Dlubal’s two analysis programs RSTAB and RFEM can now beupgraded with an add-on module for the design of aluminiummembers according to European standard 1999-1-1:2007 (EC 9).ALUMINIUM is the answer to designing for tension, com-pression, bending, shear and combined internal forces.

The program is able to take into account various cross-sections,e.g. I-sections (symmetric and unsymmetric), C-sections, T-sections,rectangular hollow sections, square sections, round sections, an-gles with equal and unequal legs, steel flats and round bars.

Features of ALUMINIUMIn addition to analysing general internal forces, ALUMINIUMdesigns the structure for stability and serviceability. The programconsiders several stability cases, such as flexural buckling, tor-sional buckling and lateral-torsional buckling. The critical buck-ling load and the elastic critical moment for general loading andsupport conditions are determined automatically by means of anintegrated FEA program (eigenvalue analysis). Furthermore, it ispossible to define lateral intermediate supports for members se-lected for the design.Serviceability limit state design is carried out for the characteris-tic, frequent or quasi permanent design situation as required.During the design process in ALUMINIUM the user has thechance to optimize particular cross-sections within the samecross-section table. Subsequently, it is possible to transfer theoptimized cross-sections to RSTAB or RFEM. A cross-sectionclassification is performed automatically by the program.

Working with ALUMINIUM The mode of operation is similar to that of other Dlubal applica-tions. Structure and loads are entered in RSTAB or RFEM, andthe design is performed in the add-on module. For aluminiumdesign it is important to note that the data specified in RSTAB/RFEM for materials, loads and load combinations must be enteredin accordance with the design concept described in the Eurocode.The RSTAB/RFEM library already contains appropriate materials.ALUMINIUM provides the parameters presented in the nationalannexes (NAs) of the following countries: Germany, Italy, CzechRepublic, Cyprus, Denmark, Ireland and Slovakia. The presetvalues for intermediate lateral supports and effective lengths can

be adjusted. Where continuous members are being designed, theuser can define individual support conditions and eccentricitiesfor each intermediate node.

Output of resultsThe results tables are clearly laid out. For example, the first tableshows the maximum design ratio with the corresponding analysisresults for each load case, group or combination that has beenanalysed. Moreover, the output includes governing internalforces as well as parts lists. The complete module data is part ofthe RSTAB/RFEM hardcopy report. The report contents and theextent of the output can be selected specifically for the individualdesigns.

Further information and demo versions: Ingenieur-Software Dlubal GmbH, Am Zellweg 2, 93464 Tiefenbach, Germany, tel. +49 (0)9673 9203-0, fax +49 (0)9673 9203-51, [email protected], www.dlubal.de

A7Steel Construction 5 (2012), No. 1

Products & Projects

ALUMINIUM provides the parameters given in the national annexes of these countries:Germany, Italy, Czech Republic, Cyprus, Denmark, Ireland and Slovakia. (© Dlubal)

Grontmij to design largest shopping mall of Scandinavia

Grontmij has been awarded the structural design of the largestshopping mall in Scandinavia; the ‘Mall of Scandinavia’. The as-signment was awarded by Peab, one of the leading constructioncompanies in the Nordic region. The total contract value is € 392million. Grontmij’s fee is over € 5 million. Grontmij’s works willrun until the end of 2014.The assignment of Grontmij includes the structural design andthe coordination of sub suppliers’ engineering work related tothe structure of the shopping mall.The ‘Mall of Scandinavia’ is one of the largest building projects inSweden. After completion it will be the biggest shopping centrein Scandinavia with approximately 100,000 m2 of retail space. Themall will include a multiplex cinema and 250 shops and restau-rants. The Mall of Scandinavia is part of a larger complex whichincludes a hotel, office spaces and dwellings. It will be locatedcentrally in the Stockholm region. The opening is scheduled forautumn 2015.Claes Ullén, Project Manager Peab for the Mall of Scandinavia.‘We received a very positive response from Grontmij. Theydemonstrated clear commitment, technical competence and creativity in the contractual framework.’ Leif Bertilsson, CEO of Grontmij Sweden: ‘It’s an honour to beinvolved in this landmark project for Sweden and Stockholm inparticular. It will be one of the most exciting building projects inSweden for the next years. Together with our Danish Grontmijcolleagues we’ll be glad to contribute. This is a good example ofour international capability delivered locally.’

Further information:Grontmij N.V., De Holle Bilt 22, 3732 HM De Bilt, The Netherlands, tel. +31 30 220 79 11, [email protected],direct contact: Grontmij N.V., Arnold Drijver, Director Corporate Communication, tel. +31 30 220 75 48, Grontmij N.V., Michèle Negen, Investor Relations Manager, tel. +31 30 220 78 31

04_SC_A04-A10_1-12_SC 09.02.12 12:13 Seite 7

Versatile floor system with improved performance

Arval, the flooring brand from ArcelorMittal for sheet steelconstruction elements, can boast a new lightweight, cost-effec-tive solution for building long-span flooring structures: theCofraplus 220 floor system. The concept, developed for carparks, permits the design of continuous slabs and can also beused in combination with various floor beam configurationsin multi-storey building applications. Besides its advantagesduring construction and in the finished state, the system is awinner due to its simple, quick assembly.

The Cofraplus 220 floor system is based on a 220 mm sheetmetal section with two webs, a ribbed crown and a cover widthof 750 mm. The thickness of the sheet metal can be varied andso the system can be finely adapted to different load-span speci-fications.Car parks and multi-storey buildings benefit from the long spanswithout propping and the possibility of introducing continuousslabs for a cost-effective design. The sheets are made of highlyanti-corrosive pre-painted sheet steel, available in a choice ofpure metal coating or a number of organic coatings with variousfeatures. The range of colours is even more extensive. The lightflooring substrate produced by the sheet metal section providesfriendly, secure surroundings; colourful accents are an aid tohorizontal and vertical orientation in the building.

Continuous slabs for cost-effective loadbearing characteristicsThe floor system is dimensioned taking into account its additiveeffect. Here, the concrete cross-section of the ribbed floor andthe sheet metal are each involved in the load transfer separately,with no connection to each other. The total loadbearing capacityof the floor is thus achieved through simple addition of the twoloadbearing portions. In addition to the considerable moment ofinertia of the 220 mm floor slab, particularly favourable in criticaldesigns, the Cofraplus 220 System permits a static continuous slabdesign above the beams acting as intermediate supports. The sheetmetal sections are supported via so-called wings: 3 mm thicksheet metal parts fitted to the profile, which are factory-weldeddirectly to the web of the composite steel beams. The attachmentof the wings to the web can vary depending on the specific proj-ect. If required, the crown of the sheet metal section can be po-sitioned over the upper edge of the support flange. In this way,the composite steel beam, as the primary loadbearing element,can be optimized in a more independent manner without in-creasing the depth of the concrete floor between beams or in-creasing the loads. Since they are supported via the wings, thesheet metal sections are brought right up to the web, along withthe ribs of the floor system which will be concreted at a laterdate. This construction achieves a high transverse loadbearingcapacity between support and sheet metal, providing the prereq-uisite for redistribution of moments of inertia and enabling amore cost-effective design on the basis of a continuous slab.


Products & Projects

Fig. 1. Concreting the Corfaplus 220 floor system

Fig. 2. Connection detail of the Cofraplus 220 floor system – welded support “wings”and flooring

Fig. 3. Cofraplus 220 floor system with continuity effect for better economy(© Arcelor Mittal)

Fig. 4. The system can be finely adapted to different load-span specifications

04_SC_A04-A10_1-12_SC 09.02.12 12:13 Seite 8

Generally no need for reinforcement in the corrugationsDepending on the load–span specifications and the fire resistancerequired, there is generally no need for reinforcement in the cor-rugations. Likewise, complicated transverse reinforcement overthe beams can be dispensed with. Since public car parks withspans of 15 to 17 m and beam spacings of 5.0 to 6.0 m do notrequire stricter fire resistance requirements, only anti-shrinkagereinforcement is necessary. Depending on the requirements forcrack control reinforcement, which are generally calculated basedon the corrosion protection and the planned surface finishes, itmay be necessary to allow for more reinforcement over the beams.Perimeter zones, ramps and other details must be considered onan individual basis.

Spans of up to 6.3 m without propping –even while construction is in progressThe loads generated in the construction phase are generally acrucial factor when designing long-span floor slabs because thefresh concrete has not yet solidified and the “permanent form-work” is the only element ensuring load transfer. Thus, it is thecondition of the concrete that represents a critical load for thesheet metal section, even if a more favourable safety factor inthe design concept can be selected here. Depending on the weightof the concrete and the depth of the floor above the sheet metal,the Cofraplus 220 System permits spans of up to 6.3 m withoutpropping in the construction phase. If temporary joists are usedduring the construction phase, a clear span of up to 9.0 m iseven possible in the final construction. In such cases, the ribs

A9Steel Construction 5 (2012), No. 1

Products & Projects


Kundenservice: Wiley-VCHBoschstraße 12D-69469 Weinheim


ww

w.e

rnst

-un

d-s

oh

n.d

e/j

ou

rna

ls

Order a free sample copy: www.ernst-und-sohn.de/structural-concrete


P u b l i s h e r : f i b – I n t e r n a t i o n a l F e d e r a t i o n f o r S t r u c t u r a l C o n c re t e

Structural ConcreteJournal of the fi b

Volume 13, 20124 issues / year.Editor: Luc Taerwe (Belgium)Deputy Editor: Steinar Helland (Norway)Journal for fi b members

Annual subscription printISSN 1464-4177for companies € 120,–*for libraries € 460,–*

Annual subscribtion print + onlineISSN 1751-7648for companies € 138,–*for libraries € 529,–*

Accredited at ISI since volume 2009.

Structural Concrete, the offi cial journal of the fi b, provides conceptual and procedural guidance in the fi eld of concrete construction, and features peer-reviewed papers, keynote research and industry news covering all aspects of the design, construction, performance in service and demolition of concrete structures. Starting in January 2011 it will be part of the journals portfolio of Berlin-based publishing house Ernst & Sohn.

*Pric

es v

alid

unt

il A

ugus

t 31

, 201

2. P

rices

are

net

-pric

es e

xcl.

of V

AT

but

incl

. pos

tage

. Pric

es a

re s

ubje

ct t

o al

tera

tions

. 01

98

10

00

06

_dp

Knowledge where you need it

must be held by appropriate supports in order to prevent a localstability failure.

Prefabrication and erection with minimum tool requirementsThe composite steel beams are already prefabricated to a largeextent, including wings. This enables extremely fast and easyerection because the system generally requires no time-consum-ing assembly of complex parts and provides a thick shell for theconcreting process. For sheet metal mounted at a higher level,openings between the steel beams and the floor sheets are sim-ply closed off with a small, continuous Z-section. Owing to theirlight weight, the sheet metal panels can easily be positioned oneach floor, saving time and crane capacity. A structure withoutpropping in the construction phase therefore makes it possibleto concrete several floors at a time without having to wait forthe concrete to harden.In conjunction with an integrated Slim Floor Beam from Arcelor-Mittal (CoSFB – Composite SLIM FLOOR BEAM), the systemcan achieve clear grid dimensions of approx. 9 × 10 m. This un-derscores the performance of the Cofraplus 220 floor system asa lightweight, cost-effective solution for long-span floor structuresin multi-storey buildings, and particularly in car parks.

Further information:ArcelorMittal Construction, International Arval, Industriestr. 19, 8112 Otelfingen/Zurich, tel. +41 (0)56 2961010, www.arcelormittal.com

04_SC_A04-A10_1-12_SC 09.02.12 12:13 Seite 9

*Pr

ices

are

val

id in

Ger

man

y, e

xclu

sive

ly, a

nd s

ubje

ct t

o al

tera

tions

. Pric

es in

cl. V

AT.

Boo

ks e

xcl.

ship

ping

. Jou

rnal

s in

cl. s

hipp

ing.

019

9100

006_

dp

On

lin

e-O

rde

r: w

ww

.ern

st-u

nd

-so

hn

.de





This book introduces the basis design concept of Eurocode 3 for current steel structures, and their practical application. Numerous worked examples will facilitate the acceptance of the code and provide for a smooth transition from earlier national codes to the Eurocode.

E D . : E C C S – E U R O P E A N

C O N V E N T I O N F O R

C O N S T R U C T I O N A L S T E E LW O R K

Design of Steel StructuresEC3: Design of Steel Structures. Part 1-1: General Rules and Rules for Buildings.

April 2010.446 pages, 295 fi g., 105 tab., Softcover.

70,–*

ISBN 978-3-433-02973-2

E D . : E C C S

Design of Cold-formed Steel StructuresEC3: Design of Steel Structures.Part 1-3: Design of Cold-formed Steel Structures.

201 . approx. 512 pages. approx. 300 fi g. Softcover.approx. 70,–*

ISBN 978-3-433-02979-4

The book is concerned with design of cold-formed steel structures in building based on the Eurocode 3 package, particularly on EN 1993-1-3. On this purpose, the book contains the essentials of theoretical background and design rules for cold-formed steel sections and sheeting, members and connections.

This volume elucidates the design rules for connections in steel and composite structures which are set out in Eurocode 3 and 4. Numerous examples illustrate the application of the respective design rule.

This design manual provides practical advice to designers of plated structures for correct and efficient application of EN 1993-1-5 design rules and includes numerous design examples.

E D . : E C C S

Design of Plated StructuresEC3: Design of Steel Structures.Part 1-5: Design of Plated Structures.

January 2011. 272 pages, 139 fi g., 19 tab., Softcover.

55,–*

ISBN 978-3-433-02980-0

E D . : E C C S

Fatigue Design of Steel and Composite StructuresEC3: Design of Steel Structures.Part 1-9: Fatigue.EC4: Design of Composite Steel and Concrete Structures.

2011. 3 pages. 250 fi g. Softcover

55,–*

ISBN 978-3-433-02981-7

This volume addresses the specific subject of fatigue, a subject not familiar to many engineers, but still relevant for proper and good design of numerous steel structures. It exp-lains all issues related to the subject: Basis of fatigue design, reliability and various verification formats, determination of stresses and stress ranges, fatigue strength, application range and limitations. It contains detailed examples of application of the concepts, computation methods and verifications.

E D . : E C C S

Fire Design of Steel StructuresEC1: Actions on Structures. Part 1-2: Actions on Structuresexposed to Fire. EC3: Design of Steel Structures. Part 1-2: Structural Fire Design.

May 2010.428 pages, 134 fi g., 21 tab., Soft-cover.

70,–*

ISBN 978-3-433-02974-9

This publication sets out the design process in a logical manner giving practical and helpful advice and easy to follow worked examples that will allow designers to exploit the benefits of the new approach given in the Eurocodes to fire design.

E D . : E C C S

Design of Connections in Steel

and Composite Structures201 .

approx. 500 pages, approx. 300 fi g.

Softcover.

approx. 70,–*

ISBN: 978-3-433-02985-5

prel

imin

ary

prel

imin

ary

04_SC_A04-A10_1-12_SC 09.02.12 12:13 Seite 10

1© Ernst & Sohn Verlag für Architektur und technische Wissenschaften GmbH & Co. KG, Berlin · Steel Construction 5 (2012), No. 1

Editorial

Large steel structures such as bridge girders or girders forheavy overhead cranes are usually built up from steelplates welded together – and are so slender that they arehighly prone to buckling. With the introduction of theEurocodes, in most European countries the design ofslender steel plates is now covered by Eurocode 3Part 1.5. Within the process of European harmonization,traditional methods as well as new research develop-ments have found their way into the code put togetherby a very knowledgeable project team headed by BerntJohansson. Earlier publications, also in this journal (e.g.“Steel Construction”, No. 4, 2009), have provided somebackground.

Although the code covers many common stabilityproblems for plated structures, there is still a need for,

and possibility of, further development. For a long time, the technical development ofrules for plate buckling has been driven forward by ECCS Technical Working Group(TWG) 8.3, a subgroup of ECCS Technical Committee (TC) 8 “Stability”. Again andagain over the years, research issues have often been motivated by questions comingfrom practice, such as bridge launches and the introduction of high concentratedforces (patch loading) associated with this, or the design of plated structures stiffenedby modern trapezoidal stiffeners with high torsional rigidity. These are some of the top-ics discussed by experts from universities and industry. Joint research has been initi-ated and has resulted in several PhD theses with supervisors from different Europeanuniversities. So true European collaboration has evolved, with some of the results be-ing presented in “Steel Construction”. An overview of recent developments, with a spe-cial focus on bridge design, was presented by Ulrike Kuhlmann, Benjamin Braun,Hervé Degée and Antonio Zizza in “Steel Construction” No. 4, 2011. A more extensivepresentation with five papers of recent results of the work of ECCS TWG 8.3 can befound in this issue.

The good collaboration between members of TWG 8.3 has also resulted in theECCS Eurocode design manual “Design of Plated Structures”. This was produced inorder to provide practical advice to designers of plated structures, so that they can ap-ply the EN 1993-1-5 design rules correctly and efficiently.

On the other hand, CEN/TC250/SC3, responsible for Eurocode 3 in general, hascreated several expert groups (called Evolution Groups) for the maintenance and on-going development of the different parts of Eurocode 3; one of those groups is the Evo-lution Group for EN 1993-1-5. As the two groups ECCS TWG 8.3 and CEN/TC250/SC3/EG 1-5 had a chairperson and many members in common, it was decided to bringboth together in order to avoid duplication of effort. The first joint meeting took placein May 2011 in Paris, a second one in Stuttgart in November 2011. The first amend-ments for the future version of EN 1993-1-5 have been discussed and are partially re-flected in the contributions on the following pages.

In this issue there is a focus on the patch loading situation, which is typical forcrane runway beams and launching of bridge girders. An economically interesting al-ternative to homogeneous girders for such applications is the hybrid steel plate girder,

Eurocode 3 Part 1.5: Plated Structures –new chances and developments

Ulrike Kuhlmann

DOI: 10.1002/stco.201210010

05_001-002_Editorial_SC 09.02.12 12:15 Seite 1

2 Steel Construction 5 (2012), No. 1

with flanges of high-strength steel and a web of mild steel. Rolando Chacón, MarinaBock, Enrique Mirambell and Esther Real report on research into patch loading resis-tance – and reveal contradictions in the existing code rules. The method they devisedis integrated into a new resistance model for steel plate girders subjected to patch load-ing, which has widely been discussed in TWG 8.3 because it summarizes the findings ofseveral doctoral theses and offers a new, improved formulation for EN 1993-1-5. Thestatistical evaluation of this new resistance model is presented by Rolando Chacón,Benjamin Braun, Ulrike Kuhlmann and Enrique Mirambell.

Another interesting design alternative to flat web plates is the corrugated web,which allows for higher shear and also patch loading forces and avoids numerous ex-pensive transverse stiffeners and frames. The joint work on the design of girders withtrapezoidal corrugated webs under patch loading, shear and bending interaction is sum-marized, also in the light of new rules for EN 1993-1-5, by Balász Kövesdi, BenjaminBraun, László Dunai and Ulrike Kuhlmann.

One of the main discussion points during the drafting of the existing Eurocode 3Part 1.5 was the design axial force in intermediate transverse stiffeners. So this subjecthas been on the agenda of TWG 8.3 for some time; it is also an important issue fromthe viewpoint of industrial practice. Meanwhile, recent research has opened up an in-teresting possibility that might solve this question through a minimum requirement forthe stiffness of the intermediate transverse stiffener. This solution and its backgroundare explained by Franc Sinur and Darko Beg.

At a very late stage of the development of EN 1993-1-5, a second method of veri-fication was introduced into the code based on reduced stresses instead of reduced “ef-fective” sections. This method is especially powerful for multi-axial stress situationsand irregular geometries for which no other design rule exists. However, its traditionsare mainly rooted in German design practice and it is not very well known in otherparts of Europe. In “Reduced stress design of plates under biaxial compression”, Ben-jamin Braun and Ulrike Kuhlmann outline the background to this method applied toa special stress state, identify deficiencies and also offer an improved solution.

At this point I want to thank all the contributors for their intensive collaborationand also the colleagues of ECCS TWG 8.3 and CEN/TC250/SC3/EG 1-5 for the manyvaluable discussions and contributions. And I would also like to thank the journal forthe opportunity of this special issue.

Prof. Dr.-Ing. Ulrike KuhlmannHead of Institute of Structural DesignUniversity of Stuttgart

Editorial

05_001-002_Editorial_SC 09.02.12 12:15 Seite 2


Hybrid girders represent an economical alternative to homoge-neous girders because they achieve greater flexural capacitywith less material. One of the potential applications of hybrid steelplate girders is their use in bridges. One potential method of con-struction for these bridges is the push launch method in whichpatch loading may affect the design. The aim of this paper is topresent the advanced conclusions of research work dealing withthese two fields simultaneously: hybrid steel plate girders sub-jected to the particular case of patch loading. It is shown that,contrary to the EN 1993-1-5 formulation, the influence of the fyf /fywratio (namely, the hybrid grade) is negligible for both unstiffenedand longitudinally stiffened girders according to the EN 1993-1-5assumptions. Suggestions for considering these findings in de-sign codes are provided at the end of the paper.

1 Introduction

The collapse behaviour of patch-loaded homogeneous plategirders has been widely reported [1, 2]. Researchers haveproposed several expressions for the elastic critical loadsand physical models which, referring to these expressions,accurately reproduce the limit state of the plates at ultimateload. Most of these approaches agree with a vast numberof experimental results obtained from various sources. How-ever, it has been pointed out that the vast majority of thesestudies only deal with the resistance of homogeneous plategirders [3–6].

The resistance of both unstiffened and longitudinallystiffened hybrid steel plate girders subjected to patch load-ing is studied in this work. The paper is based on the con-clusions previously given in [3–6]. Such works demonstratethat, contrary to the EN 1993-1-5 provisions, the varyingyield strength of the flange fyf (and hence the hybrid grade)does not play any role in the resistance of girders sub-jected to concentrated loading. Highlights concerning thisproposal are presented at the end of the paper.

2 Review of earlier work

A steel girder is deemed as being hybrid when it is fabri-cated with different steel strengths for the flange and webpanels. This type of girder is popular because compared

with a homogeneous girder, greater flexural capacity is ob-tained at lower cost and weight [7–8]. Extensive experi-mental, theoretical and numerical research on hybrid de-sign can be found in the literature [9–13].

Patch loading phenomena has been widely analysedsince the early 1960s. Experimental and theoretical analy-ses have pinpointed the failure mechanisms of girders sub-jected to patch loading and, consequently, ultimate loadpredictions have been provided [14–20].

Despite the vast amount of research devoted to hy-brid girders and patch loading separately, research workcovering both subjects is rather scant. Schilling [21] pre-sented the first publication related to hybrid steel girdersdealing explicitly with concentrated loading. Moreover,when studying the behaviour of slender girders subjectedto patch loading, Granath [22] reached conclusions aboutthe influence of the moment capacity of the flanges onthe bearing capacity of plate girders subjected to concen-trated loads. Following this thread, the authors of thispaper subsequently presented several research works[3–5].

3 EN 1993-1-5

Verification of patch loading according to EN 1993-1-5FRd is based on simplifications of the procedures providedin [1] and [2]. The general approach currently included inEN 1993-1-5 is based on a plastic resistance Fy which ispartially reduced by means of the resistance function(Eq. (1)). The plastic resistance includes the key length para-meter ly, which is the yield-prone effectively loaded length.This length can be calculated from the geometrical andmechanical properties of the girders using Eqs. (2) and (3).More details concerning the EN 1993-1-5 formulation aregiven in a companion paper [23].

(1)

(2)

(3)

FF f l t f a t

RdF y

M

F yw y w

M

F yw w

M

= = ≤χγ

χγ

χγ

· · · · · · ·

1 1 1

l s t m m ay s f= + + +( ) ≤2 1 1 2· ·

mf b

f tm

ht

yf f

yw w

w

f1 2

2

0 02= =⎛

⎝⎜⎞

⎠⎟·

·. ·

Hybrid steel plate girders subjected topatch loading

Rolando Chacón*Marina BockEnrique MirambellEsther Real

Articles

DOI: 10.1002/stco.201200001

* Corresponding author: e-mail [email protected]

06_003-009_Chacon (SC2)_000-000_Steel Constr. (2sp).qxd 09.02.12 12:21 Seite 3

4 Numerical study

A numerical study was carried out using the FE-basedsoftware Abaqus-Simulia [24]. The software was used as asimulation tool and therefore numerical databases of un-stiffened and longitudinally stiffened steel plate girderssubjected to patch loading could be modelled and con-structed. Details of those models can be found in [6].

4.1 Unstiffened girders

The simulations were performed on a single panel centri-cally loaded with a patch as sketched in Fig. 1a. The nu-merical database was constructed by varying the parame-ters depicted in Table 1, which resulted in 192 specimens.

4.2 Longitudinally stiffened girders4.2.1 Open stiffeners

The simulations were performed on a single panel centri-cally loaded with a patch as sketched in Fig. 1b. The nu-merical database was constructed by varying the parame-ters given in Table 2. Other parameters such as the distancebetween transverse stiffeners (a = 9000 mm), the web yieldstrength (fyw = 235 N/mm2), the web depth (hw = 3000 mm),the flange dimensions (80 × 900 mm), the transverse stiff-ener dimensions tst · bf (30 × 900 mm), the longitudinalstiffener width (bsl = 300 mm) and the stiff bearing length(ss = 2250 mm) were kept constant. Table 2 summarizesthe set of variations.

4.2.2 Trapezoidal closed stiffeners

Several girders were provided with a closed, trapezoidallongitudinal stiffener (Fig. 1c). In this case, both the posi-tion and the stiffener rigidity were kept constant. Otherparameters such as the distance between transverse stiff-eners (a = 9000 mm), the web yield strength fyw =


R. Chacón/M. Bock/E. Mirambell/E. Real · Hybrid steel plate girders subjected to patch loading

Fig. 1. Geometry of the modelled prototypes

Table 1. Numerical database of unstiffened girders

Table 2. Numerical database of longitudinally stiffenedgirders, open stiffeners

235 N/mm2, the web depth hw = 3000 mm, the flange di-mensions (80 × 900 mm), the transverse stiffener dimen-sions (30 × 900 mm) and the stiff bearing length ss =2250 mm were kept constant as well. Table 3 gives thegeometrical data of the specimens tested numerically. It isworth pointing out that for all cases, the longitudinal stiff-ener has the same thickness tsl as the web plate tw. Fig. 2shows the geometrical proportions of the closed longitudi-nal stiffener.

Numerical data-base variations

Group

U0 UI UII UIII

Web yieldstrength fyw(N/mm2)

235 235 235 235

Flange yieldstrength fyf(N/mm2)

235 235 235 235

275 275 275 275

355 355 355 355

460 460 460 460

hw (mm) 1000 2000 3000 4000

a (mm)

1000 2000 3000 4000

2000 4000 6000 8000

3000 6000 9000 12000

tw (mm)8 12 15 15

12 20 25 30

ss (mm)250 500 750 1000

500 1000 1500 2000

Flange dimen-sions (mm2)

800 × 60 900 × 80 1000 × 80 1200 × 100

Stiffener thick-ness (mm)

40 60 60 80

Girders pergroup

48 48 48 48

Total number of numerical simulations

192

Dimensions LS0 LSI LSII LSIII LSIV

hw (mm) 3000 3000 3000 3000 3000

tw (mm) 10 15 20 25 30

fyf (N/mm2)

235 235 235 235 235

355 355 355 355 355

460 460 460 460 460

b1 (mm)

300 300 300 300 300

600 600 600 600 600

900 900 900 900 900

1200 1200 1200 1200 1200

tsl (mm)

10 10 10 10 10

20 20 20 20 20

30 30 30 30 30

40 40 40 40 40


5Steel Construction 5 (2012), No. 1

5 Numerical results

Results in the form of ultimate load capacity Fu, load-dis-placement plots, load-stress plots and thorough compar-isons with EN 1993-1-5 are widely available in previouslypublished papers from the authors [3–5]. Tables with pre-cise data concerning each numerically simulated proto-type as well as several plots and contours of the simula-tions are also provided. In this paper, and for the sake ofconciseness, only the key results are displayed. These re-sults aim at show the most remarkable findings of the pre-sent research work.

5.1 Ultimate load capacity

The ultimate load capacity of the steel plate girders sub-jected to patch loading has been obtained as the maxi-mum load the prototype is able to carry in an incrementalprocess. The observed failure modes are in accordance withthe typical failure modes associated with patch loading.Slender girders tend to fail by local instability, whereasstocky girders tend to fail by an intertwined local yielding-buckling mode. Fig. 3 shows a qualitative plot in which lo-cal folding of the loaded panel is noticeable.

Moreover, numerical results for Fu are illustratedgraphically in Fig. 4a for unstiffened girders and Fig. 4bfor longitudinally stiffened girders. These results are plot-ted on the χ-λ space for comparison purposes.

It is possible to draw the following conclusions fromthese plots:

– All results are located above the EN 1993-1-5 (on thesafe side).

– Generally, the results follow a typical Eulerian hyper-bolic shape.

– Some girders highlighted in Fig. 4a show anomalous re-sults. Their geometrical proportions are such that theEN 1993-1-5 formulation must be modified slightly andtherefore these girders do not follow the same trends.


Fig. 3. Typical failure modes

Fig. 4. Ultimate load capacity of the numerical prototypes

Fig. 2. Geometric pro-portions of the closedlongitudinal stiffener

Table 3. Numerical database of longitudinally stiffenedgirders, closed stiffeners

Dimensions LSa LSb LSc LSd LSe

hw (mm) 3000 3000 3000 3000 3000

tsl = tw (mm) 8 10 12 15 20

fyf (N/mm2)

235 235 235 235 235

355 355 355 355 355

460 460 460 460 460

b1 (mm) 900 900 900 900 900

Web folding of the loaded panel

Web folding of both loaded panels


These girders have closely spaced transverse stiffeners(in these cases, the effectively loaded length ly is greaterthan the distance between transverse stiffeners a, i.e. ly > a) and their failure modes do not fully match withpatch loading. Additional research concerning this topichas been addressed by the authors [25].

5.2 Load-deflection plots

Load-deflection plots characterize the response of the steelplate girders subjected to patch loading. In a hypotheticalincremental process, the applied load is plotted against thevertical displacement of a point located on the loadedflange. The response is initially linear but becomes gradu-ally non-linear until the load reaches a maximum. Fromthis maximum load onwards, a softening branch is recordedand the deformation increases sharply.

Fig. 5 shows four load-deflection plots of selectedprototypes from the numerical database of unstiffened


girders. The main varying parameter between the four plotsis the web depth of the specimens and therefore the webslenderness hw/tw. Fig. 5a represents the stockiest girders,whereas Fig. 5d represents the most slender prototypes.

Within each plot, load-displacement curves as well asvalues of Fu are shown for girders with a varying hybridgrade fyf/fyw. The following features can be gleaned fromthese plots:– Irrespective of the web slenderness, all load-displace-

ment curves are identical because the fyf/fyw ratio is var-ied. The response of the girders does not depend on thevariation of the flange yield strength.

– Irrespective of the web slenderness, the ultimate loadcapacities Fu are identical because the fyf/fyw ratio isvaried. Fu is not dependent on the flange yield strengtheither.

– Stocky girders present a quite marked linear branch,whereas in slender prototypes the curves are non-linearfrom low levels of load onwards.


Fig. 5. Load displacement plots for unstiffened girders



On the other hand, Fig. 6 shows four load-deflection plotsof selected prototypes from the numerical database of lon-gitudinally stiffened girders. The main varying parameterbetween the four plots is the position of the longitudinalstiffened b1 and therefore the b1/hw ratio. Fig. 6a repre-sents the longitudinal stiffener closest to the loaded flange(b1/hw = 0.1), whereas Fig. 6d represents the longitudinalstiffener with a relative position b1/hw = 0.4.

Within each plot, load-displacement curves as well asvalues of Fu are shown for girders with varying hybridgrade fyf/fyw. The following features should be pointedout:– Irrespective of the position of the longitudinal stiff-

ener, all load-displacement curves are identical be-cause the fyf/fyw ratio is varied. The response of thegirders does not depend on the variation of the flangeyield strength.

– Irrespective of the position of the longitudinal stiffener,the ultimate load capacities Fu are identical because thefyf/fyw ratio is varied. Fu is not dependent on the flangeyield strength either.

5.3 Hybrid vs. homogeneous prototypes – numerical andEN 1993-1-5 results

The ultimate load capacity Fu of a given hybrid girder Fu,hybcan be compared with the ultimate load capacity Fu of itsequivalent homogeneous girder Fu,hom. If Fu depends onfyf/fyw, then Fu,hyb should differ from Fu,hom. Fig. 7 showsthe ratio Fu,hyb/Fu,hom as fyf/fyw is varied for unstiffenedgirders (Fig. 7a) and longitudinally stiffened girders (Fig. 7b).Numerical as well as EN 1993-1-5 results are included ineach plot.

From these figures, the following features should bepointed out:– For both unstiffened and longitudinally stiffened gird-

ers, the numerical results show very little dependencyon fyf/fyw. The Fu,hyb/Fu,hom ratios vary, ranging fromFu,hyb/Fu,hom = 1.00 to 1.02 (variation that might ar-guably be attributed to numerical reasons).

– For both unstiffened and longitudinally stiffened gird-ers, the EN 1993-1-5 results show a significantly strongdependency on fyf/fyw. The Fu,hyb/Fu,hom ratios vary,


Fig. 6. Load displacement plots for longitudinally stiffened girders


ranging from Fu,hyb/Fu,hom = 1.02 to 1.14. This anomalyis structurally unsafe and should be corrected in the for-mulation.

6 Design proposal

The results obtained suggest that the patch loading resis-tance should not depend on the hybrid parameter fyf/fywto any extent because the flange yield resistance does notseem to play any role in the development of the collapsemechanism.

For the sake of correcting the aforementioned anom-aly, a modification to the current EN 1993-1-5 formula-tion that enhances the results quite satisfactorily is pro-vided. Readably, it is proposed that the m1 coefficient inEq. (3) be replaced by m1

*, see Eq. (4):

(4)

This proposal has been tested both structurally and sta-tistically. The results lead to a satisfactory improvementof the formulation. The results obtained with the up-graded coefficient m1

* are structurally sound and on thesafe side.

l s tbt

ht

mb

y s ff

w

w

f

f= + + +⎛

⎝⎜⎞

⎠⎟

⎛

⎝

⎜⎜

⎞

⎠

⎟⎟ =2 1 0 02

2

1· , · *

ttw


7 Conclusions

A vast numerical study of more than 800 hybrid and ho-mogeneous steel plate girders has been presented in thispaper. The numerical study is based on girders with geo-metrically realistic proportions. The parametric variationof the most relevant variables influencing the resistance ofgirders to patch loading is quite profuse, although the par-ticular emphasis was on the hybrid grade fyf/fyw.

The numerical results presented for both transver-sally and longitudinally stiffened girders do not agree withthe results provided by EN 1993-1-5 when the focus is onthe effect of the flange yield strength (and therefore thehybrid parameter fyf/fyw). Numerically, it is predicted thatfor girders with widely spaced transverse stiffeners (ly < a),fyf/fyw has no influence on the ultimate load capacity ofpatch-loaded girders. The current formulation of EN 1993-1-5 takes this ratio into account in such a way that thegreater the ratio fyf/fyw, the higher is the ultimate load ca-pacity of the girders. It is worth pointing out that this pa-rameter appears explicitly in the term m1.

The current formulation of EN 1993-1-5 leads to struc-turally unsafe results. An attempt to correct this anomalyinvolves making the fyf/fyw ratio equal to 1.0 in the currentexpressions for FRd. This attempt leads to reshaping the effectively loaded length ly in Eq. (4), in which m1 is re-placed by m1

*. This proposal has been tested both struc-turally and statistically in [3–5]. For reliability purposes,further research concerning the recalibration of the resis-tance function and the partial safety factor is addressed ina companion paper [23].

Acknowledgements

The authors gratefully acknowledge the financial supportprovided by the Ministerio de Ciencia e Innovación as partof research project BIA-2008-01897 as well as the scholar-ship provided by AGAUR to Ms Bock.

References

[1] Lagerqvist, O., Johansson, B.: Resistance of I-girders to con-centrated loads. Journal of Constructional Steel Research,vol. 39, 1996, pp. 87–119.

[2] Graciano, C.: Patch Loading: Resistance of longitudinallystiffened steel girder webs. Doctoral Thesis 2002:18, LuleaUniversity of Technology, Sweden. ISRN: LTU-DT-02/18-SE,2002.

[3] Chacón, R., Mirambell, E., Real, E.: Hybrid steel plate gird-ers subjected to patch loading. Part 1: Numerical study. Jour-nal of Constructional Steel Research, vol. 66 (5), 2010,pp. 695–708.

[4] Chacón, R., Mirambell, E., Real, E.: Hybrid steel plate gird-ers subjected to patch loading. Part 2: Design proposal. Jour-nal of Constructional Steel Research, vol. 66 (5), 2010,pp. 709–715.

[5] Chacón, R., Bock, M., Real, E.: Longitudinally stiffened hybrid steel plate girders subjected to patch loading. Journalof Constructional Steel Research, vol. 67 (9), 2011,pp. 1310–1324.

[6] Chacón, R., Mirambell, E., Real, E.: Influence of designer-assumed initial conditions of the numerical modelling of steelplate girders subjected to patch loading. Thin-Walled Struc-tures, vol. 47 (4), 2009, pp. 391–402.


Fig. 7. Fu,hyb/Fu,hom



[7] Barker, M. G., Schrage, S. D.: High performance Steel BridgeDesign and Cost Comparisons. Transportation ResearchRecord. No. 1740, 2000, pp. 33–39.

[8] Veljkovic, M., Johansson, B.: Design of hybrid steel girders.Journal of Constructional Steel Research, vol. 64, 2004,pp. 535–547.

[9] Greco, N., Earls, C.: Structural Ductility in Hybrid HighPerformance Steel Beams. Journal of Structural Engineering,ASCE, vol. 129, No. 12, 2000.

[10] Carskaddan, P.: Shear buckling of unstiffened hybrid beams.Journal of the structural division, ASCE, vol. 94 (ST8), 1968,pp. 1965–1990.

[11] Nethercot, D.: Buckling of welded hybrid steel I-beams.Journal of the structural division, ASCE, vol. 102 (ST3), 1976,pp. 461–474.

[12] Bitar, D., Lukiç, M., Trumph, H., Galea, Y.: Compositebridge design for small and medium spans. Final report. Insti-tute of Steel Construction, ECSC Steel RTD program, 2001.

[13] Frost, R., Schilling, C.: Behaviour of hybrid beams sub-jected to static loads. Journal of the Structural Division ASCE,vol. 90 (ST3), 1964, pp. 55–86.

[14] Elgaaly, M.: Web design under Compressive Edge loads.Engineering Journal, vol. 20 (4), 1983, pp. 153–171.

[15] Roberts, T., Rockey, K.: A mechanism solution for predict-ing the collapse loads of slender plate girders when subjectedto in-plane patch loading. Proceedings of the Institution ofCivil Engineers, Part 2, vol. 67, 1979, pp. 155–175.

[16] Kutmanova, I., Skaloud, M.: Ultimate Limit State of Slen-der Steel Webs Subject to (i) Constant and (ii) Repeated Par-tial Edge Loading. Journal of Construction Steel Research,vol. 21, 1992, pp. 147–162.

[17] Markovic, N., Hajdin, N. A.: Contribution to the analysisof the behaviour of plate girders subjected to patch loading.Journal of Constructional Steel Research, vol. 21, 1992,pp. 163–173.

[18] Graciano, C., Casanova, E.: Ultimate strength of longitudi-nally stiffened I-girder webs subjected to combined patch load-

ing and bending. Journal of Constructional Steel Research,vol. 61, 2005, pp. 93–111.

[19] Davaine, L., Aribert, J.: Launching of steel girder bridge –Patch load resistance of longitudinally stiffened webs. Pro-ceedings of 4th European Conference on Steel & CompositeStructures, Maastricht, The Netherlands, 8–10 Jun 2005.

[20] Seitz, M., Kuhlmann, U.: Longitudinally Stiffened GirderWebs Subjected To Patch Loading. Proceedings of Steelbridge,Millau, 23–25 Jun 2004.

[21] Schilling, C.: Web Crippling Test on Hybrid Beams. Journalof the Structural Division, ASCE, vol. 93, 1967.

[22] Granath, P.: Behaviour of slender plate girders subjectedto patch loading. Journal of Constructional Steel Research,vol. 42, 1997, pp. 1–19.

[23] Chacón, R., Braun, B., Kuhlmann, U., Mirambell, E.: Sta-tistical evaluation of the new resistance model for steel plategirders subjected to patch loading. Steel Construction, vol. 5,Jan 2012.

[24] Abaqus Inc, v. 6.11. Simulia, Dassault Systemes, 2011.[25] Chacón, R., Mirambell, E., Real, E.: A mechanism solution

for predicting the collapse loads of girders subjected to patchloading. Proceedings of 5th European Conference on Steel &Composite Structures, Graz, Austria, pp. 159–164, 3–5 Sept2008, ISBN 92-0147-00-90.

Keywords: hybrid girders; patch loading

Authors:Dipl.-Ing. Marina Bock Dr.-Ing. Rolando Chacón Prof. Dr.-Ing Enrique Mirambell Assoc. Prof. Dr.-Ing Esther Real Construction Engineering Department, Universitat Politècnica de Catalunya,C/Jordi Girona (Campus Nord), 1-3, C1 Building, 08034 Barcelona, Spain


The other members of the Board of Ma-nagement are Dr. Karlheinz Blessing, ChiefExecutive Officer with responsibility for theCommercial Division, Fred Metzken, Execu-tive Officer, Financial Division and PeterSchweda, Executive Officer, Human Re-sources and Labour director.

Dr. Norbert Bannenberg reappointed to the Board of Management

The Supervisory Board of the Aktien-Ge-sellschaft der Dillinger Hüttenwerke hasagain appointed Dr. Norbert Bannenberg(59) as a member of the Board of Manage-ment for the period 1 June 1 to 31 Decem-ber 2015.

Dr. Bannenberg has been a member ofthe Dillinger Hütte Board of Managementsince 1 June 2002, and is responsible forthe „Technology“ division.

Peoples

Norbert Bannenberg reappointed to the Board of Management until 31 December 2015


10 © Ernst & Sohn Verlag für Architektur und technische Wissenschaften GmbH & Co. KG, Berlin · Steel Construction 5 (2012), No. 1

Articles

This paper presents a new resistance model for steel plate gird-ers subjected to patch loading. The model aims to correct theEN 1993-1-5 formulation concerning the resistance to transverseforces with the help of findings presented in several doctoralstudies on this topic. In addition, a statistical calibration of thenew model is addressed. This calibration includes a new pro-posal for the resistance function χ-λ. The proposal is conceivedfor fulfilling safety levels of the partial safety factor that matchthe current values included in EN 1993 for instability-related prob-lems. These new proposals are in accordance with the designphilosophy found in EN 1993-1-5 for the design of plated struc-tures.

1 Introduction

The adequate and safe formulation of the patch loadingresistance is crucial in the design of launched steel bridges.The resistance of plate girders subjected to patch loadinghas been studied thoroughly. A large number of contribu-tions concerning the web resistance to patch loading areavailable in the literature. Two milestone research works[1, 2] provide the framework for the formulation for thepatch loading resistance presently implemented withinEN 1993-1-5 [3]. In these works, the patch loading resis-tance is formulated by following a similar procedure foundin other instability-based verifications, namely the χ-λ ap-proach. In this approach, plastic and elastic critical buck-ling resistances (Fy and Fcr respectively) are defined sepa-rately and subsequently related to the ultimate load by aresistance function χ = f(λ).

Recent contributions [4–9] underpin the need for fur-ther modification of the plastic resistance Fy which ap-pears in the current version of EN 1993-1-5. According tothese studies, the definition of the plastic resistance over-estimates the patch loading capacity in certain cases (hy-brid girders), whereas this capacity is slightly underesti-mated for others (very slender girders).

This paper provides a brief description of the modifi-cation needed. Furthermore, a design proposal includinga new patch loading resistance model and a new resis-tance function – calibrated based on hundreds of experi-mental and numerical results – are outlined.

2 Current formulation in EN 1993-1-5

The verification of patch loading FRd is based on simplifi-cations of the procedures provided in [1] and [2]. The gen-eral approach currently included in EN 1993-1-5 is basedon a plastic resistance Fy which is reduced by means ofthe resistance function (Eq. (1)). The plastic resistance in-cludes the length ly, which can be calculated from the geo-metrical and mechanical properties of the girders usingEqs. (2) and (3). Plate instability is accounted for by meansof Eq. (4). The buckling coefficient kF is given in Eqs. (5),(6) and (7). This formulation can be applied even if thepatch load is applied on only one side of the steel girderwith the assumption that the flange is prevented from ro-tating about its longitudinal axis. The plastic resistance aswell as the elastic critical buckling loads are both relatedto the ultimate load by means of the resistance function(Eq. (8)) and the relative slenderness (Eq. (9)). The formu-lation includes a partial safety factor γM1.

(1)

(2)

(3)

(4)

(5)

(6)

(7)

(8)

FF f l t f a t

RdF y

M

F yw y w

M

F yw w

M

= = ≤χγ

χγ

χγ

· · · · · · ·

1 1 1

l s t m m ay s f= + + +( ) ≤2 1 1 2· ·

mf b

f tm

ht

if othe

yf f

yw w

w

f

F

1 2

2

0 02

0 5

= =⎛

⎝⎜⎞

⎠⎟

≥

·

·. ·

. ,λ rrwise m2 0=

F k Ethcr F

w

w

= 0 93

. · · ·

khaFw= +

⎛

⎝⎜⎞

⎠⎟6 2

2

·

kha

baF

ws= +

⎛

⎝⎜⎞

⎠⎟+ −

⎛

⎝⎜⎞

⎠⎟6 2 5 44 0 21

2

1· . . γ

γ ssl

w w w

I

h ta

hba

= ≤⎛

⎝⎜⎞

⎠⎟+ −

⎛

⎝⎜⎞

⎠10 9 13 210 0 3

3

3

1.·

· · . ⎟⎟

χλF

F

= ≤0 5 1 0. .

Statistical evaluation of the new resistance model forsteel plate girders subjected to patch loading

Rolando Chacón*Benjamin BraunUlrike KuhlmannEnrique Mirambell

DOI: 10.1002/stco.201200002



(9)

3 Recent discussions on the EN 1993-1-5 formulation

Ever since the implementation of the latest version ofEN 1993-1-5, major studies concerning the patch loadingresistance of both homogeneous and hybrid transversallyand longitudinally stiffened girders have been undertakenin France [4], Sweden [5], [6] and Spain [7–9]. These stud-ies have pointed out that, due to safety concerns, the ef-fectively loaded length ly given in Eq. (2) requires a majormodification.

3.1 Research work performed in France

In 2005, Davaine [4] studied the resistance of longitudi-nally stiffened steel plate girders subjected to patch load-ing. The main purpose of her thesis was to validate theaforementioned χ-λ approach. The question was whetherthe χ-λ approach was valid or not for the domain of steelplate girders with realistic proportions and not only thosetested in laboratories worldwide (the main basis used by[1] and [2] to calibrate their models). After a rigorousanalysis based upon results obtained with the FE models,it was suggested to maintain the χ-λ approach for the re-sistance to patch loading. Despite the fact that the resultswere deemed to be rather too far on the safe side, this ap-proach led to the best statistical fit when compared withother methods. A new resistance function for recalibratingEN 1993-1-5 was provided based on the work presentedby Müller [10]. Nevertheless, a very important issue con-cerning the patch loading phenomenon arose: throughoutthe development of the thesis, Davaine questioned thephysical meaning of the term m2 (see Eq. (3)). The thesisdid not focus on understanding the validity of this term,but it paved the way for further studies.

3.2 Research work performed in Sweden

Subsequently, major doctoral studies [5, 6] were conductedin Sweden. These studies followed the trends depicted by[1], [2] and [4] concerning the validity of the χ-λ approachas the fundamental method for calculating the resistanceof steel plate girders subjected to patch loading. Numeri-cal studies on steel plate girders subjected to patch load-

λFyw y w

cr

f l t

F=

· ·


ing were carried out for the sake of evaluating the validityof the prediction of the effective loaded length ly (Eq. (2)).In particular, the main focus was on the term m2. The ef-fective loaded length was determined from the numericalobservations according to Eq. (10). Such works drew sim-ilar conclusions to those in [4].

3.3 Research work performed in Spain

Chacón et al. [7–9] have presented a research work onpatch loading based on a larger research project concern-ing the design of hybrid girders (with fyf > fyw). The nu-merical results presented for both transversally and longi-tudinally stiffened girders did not agree with the resultsprovided by EN 1993-1-5 when the focus was on the effectof the flange yield strength (and hence the hybrid parameterfyf/fyw). Numerically, it was predicted that fyf/fyw has ab-solutely no influence on the ultimate load capacity of patch-loaded girders. The current formulation in EN 1993-1-5takes this ratio into account in such a way that the greaterthe ratio fyf/fyw, the higher is the ultimate load capacity ofthe girders. It is worth pointing out that this parameter ap-pears explicitly in the term m1.

Summarizing, Fig. 1 shows the timeline of studiesconcerning the patch loading resistance in recent years. Inaddition, although not referenced in the paper, the figurelists the researchers who set up the framework for the re-sistance of steel plate girders subjected to patch loadingfrom 1960 onwards.

4 New proposals for patch loading resistance

Two independent modifications of the patch loading resis-tance have been separately proposed to ECCS TechnicalWorking Group TWG.8.3 (Plate Buckling).

4.1 Deletion of m2

The results obtained by Davaine [4], Gozzi [5] and Clarin[6] consistently show a better prediction of the patch load-ing resistance when the term m2 is omitted from the for-mula. According to these studies, the effective loaded lengthshould be modified (Eq. (10)):

(10)l s tf b

f ty s fyf f

yw w

= + +⎛

⎝⎜⎜

⎞

⎠⎟⎟

2 1··

·

R. Chacón/B. Braun/U. Kuhlmann/E. Mirambell · Statistical evaluation of the new resistance model for steel plate girders subjected to patch loading

Fig. 1. Timeline of recent studies concerning patch loading




4.2 Replacing m1 with m1*

The results obtained by Chacón et al. [7–9] suggest thatthe patch loading resistance should not depend on the hy-brid parameter fyf/fyw to any extent because the flange yieldresistance does not seem to play any role in the develop-ment of the collapse mechanism.

For the sake of correcting the aforementioned anomaly,a modification of the current EN 1993-1-5 formulationprovides quite satisfactory results. Thus, it is proposed, asgiven below, that the m1 coefficient be replaced by m1

*

(Eq. (11)):

(11)

4.3 Final modification

After scrutiny and debate within TWG 8.3, it has been decided to merge both proposals into a single formula(Eq. (12)) that predicts the effective loaded length when asteel plate girder carries a concentrated load. This finalproposal is based on identical assumptions of the priorformulation.

(12)

For the sake of completeness, a recalibration of the resis-tance function χ-λ was performed in a collaborative taskbetween the research group from the University of Stuttgart[11] and the group from the UPC [12].

5 Resistance function for patch loading χF-λF

The resistance function relates the reduction factor χF tothe relative slenderness λF of the plate. The current formu-lation χF-λF for the patch loading resistance according toEN 1993-1-5 is based on slight modifications of the pro-posals presented in [1] and [2] (Eq. (13)):

(13)

This resistance function has been statistically calibrated tofit both the plastic resistance Fy and the elastic critical buck-ling load Fcr depicted in Eqs. (1) to (9) and to merge bothmagnitudes in a single formulation. The calibration has beenundertaken primarily with the experimental databases ex-isting until 1995 for unstiffened girders and 2002 for lon-gitudinally stiffened girders.

It is understood that any major change to such a for-mula implies recalibrating the resistance function. Fur-thermore, where additional experimental or numericalstudies are available, these prototypes may be included inthe pool of tests to obtain a higher level of statistical sig-nificance. For the sake of recalibrating the resistancefunctions with the newly proposed plastic resistance Fy,two diploma theses have been recently performed at theUniversity of Stuttgart and at the Technical University ofCatalonia [11, 12]. These works gather together all the ex-perimental and numerical tests available on the welded

χλF

F= 0 5.

l s tbt

hty s f

f

w

w

f

= + + +⎛

⎝⎜⎞

⎠⎟

⎛

⎝

⎜⎜

⎞

⎠

⎟⎟2 1 0 02

2

· . ·

l s tbty s f

f

w

= + +⎛

⎝⎜⎜

⎞

⎠⎟⎟

2 1·

steel plate girders subjected to patch loading. These data-bases include all the specimens used by [1] and [2] fortheir prior calibration as well as the newer tests availablesince then.

After consensus within TWG 8.3, the group suggestedthat any modification of the resistance curve should bebased on the proposal presented by Müller [10], whichtakes the form of the equation included in EN 1993-1-5,Annex B. This proposal is advantageous since it harmo-nizes the resistance function shape for all verifications ofcompressed members. Eqs. (14) and (15) show the generalform of this proposal:

(14)

(15)

In these equations, αF0 and λ–F0 are tuneable magnitudesthat may be calibrated for the sake of achieving the de-sired level of safety. The former is an imperfection factorand the latter the plateau length of the resistance function.For the sake of obtaining sound values for these magni-tudes, a systematic study has been carried out to establishthe sensitivity of the safety level of the formulation basedon these values.

The main random variable accounted for in this sta-tistical study is the ratio between the ultimate load capac-ity Fu and the EN 1993-1-5 prediction FRk (b = Fu/FRk).

The main statistics used in the sensitivity evaluationare the mean, the minimum and maximum values bmin andbmax, the standard deviation, the coefficient of variationand the percentage of unsafe values of the random variable(if bi < 1.0, it is not on the safe side).

Table 1 shows the sensitivity analysis of the afore-mentioned magnitudes when both αF0 and λ–F0 are system-atically varied in unstiffened girders. Table 2 shows a sim-ilar table for longitudinally stiffened girders. Different rowsrepresent different values of the plateau length λ–F0[0.2;0.3; 0.4; 0.5; 0.6; 0.7], whereas different columns representdifferent values of the imperfection factor αF0[0.25; 0.50;0.75; 1.0].

Both tables employ colours for a better understand-ing. The sensitivity of the formulation of λ–F0 can be ob-served by following a given column from top to bottom.Closer inspection leads to the following conclusions:– The greater the plateau length, the less conservative is

the formulation. The mean, minimum and maximumvalues decrease, whereas the percentage of values be-low the safety threshold increases.

– The standard deviation and the variation remain nearlyconstant, which is statistically sound.

The sensitivity of the formulation with regard to the im-perfection factor αF0 can be assessed by following columnswith the same colours (e.g. all red columns). Closer in-spection of the table leads to the following conclusions:– The greater the imperfection factor, the safer is the for-

mulation.– The coefficient of variation remains constant, whereas

the standard deviation increases slightly.

χφ φ λ

F

F F F

=+ −

≤1 01 0

2

,,

φ α λ λ λF F F F F= + −( ) +( )12

1 0 0·


It can be seen that any decision concerning the imperfec-tion factor and the plateau length alters the statistics re-lated to the safety of the formulation. As the formulationalso includes a partial safety factor γM1, it is inferred thatthis decision would also alter the value of γM1 in the fol-lowing way:– If a combination of αF0 and λ–F0 is selected from the top

left area of Tables 4 and 5 (a less safe combination), thepartial safety factor would be rather high (> γM1 = 1.1)because it would “correct” the prior, unsafe level.

– If a combination of αF0 and λ–F0 is selected from the bot-tom right area of Tables 4 and 5 (a very safe combina-tion), the partial safety factor would be rather low (tend-ing towards γM1 = 1.0).

6 Partial safety factor γM1

The partial safety factor is determined by means of thestandard evaluation procedure design method assisted bytesting described in EN 1990 [13]. The partial safety factoraccounts for the unavoidable uncertainties inherent in thedesign model, which is given by Eq. (16):

FRk = χF · fyw · ly · tw (16)

The design model comprises physical magnitudes related tothe geometry of the prototypes (width, thickness) as wellas the material of the plate (yield strength). These magni-tudes present a certain variation that has been quantifiedby means of extensive statistical studies performed mainlyby the steel industry. The procedure is based on a set of se-quential steps that involve complex statistical concepts.These steps are summarized in Table 3, together with the


main objectives and quantities involved for each one. Fur-ther information concerning this matter can be found in[11–13].

Once the procedure is applied, the values of the cor-rected partial safety factor γ*

M1 can be obtained. Studiesconcerning the sensitivity of γ*

M1 with regard to differentcombinations of αF0 and λ–F0 are available in [11] and [12].

It has been decided within TWG 8.3 to provide twodifferent combinations of the imperfection factor αF0 andthe plateau length λ–F0. The aim of these combinations is toobtain alternative values of γ*

M1 = 1.0 or γ*M1 = 1.1. As a re-

sult, once the decision is taken as to which final partialsafety factor will be used, adequate values of αF0 and λ–F0can be selected. Table 4 shows the required and adequatevalues of αF0 and λ–F0 for target values of the final partialsafety factor γ*

M1 = 1.0 or γ*M1 = 1.1.

7 Conclusions

Based on the doctoral studies presented in section 2, twoindependent modifications of the patch loading resistancehave been separately proposed to ECCS Technical Work-ing Group TWG.8.3 (Plate Buckling):– Deletion of m2 (original proposal: France and Sweden)– Replacing m1 with m1

* (original proposal: Spain)

After scrutiny and debate within TWG 8.3, it has been de-cided to merge both proposals into a single formula thatpredicts the effective length loaded when a steel plategirder carries a concentrated load. This final proposal isbased on identical assumptions of the prior formulation.For the sake of completeness, a recalibration of the resis-tance function χ-λ is depicted throughout the document.


αF0

λ–F0

0.25 0.5 0.75 1.0

0.2 1.21 0.77 2.08 16.53 0.23 0.19 1.46 0.94 2.42 2.09 0.28 0.19 1.70 1.09 2.83 0.00 0.32 0.19 1.92 1.23 3.22 0.00 0.36 0.19

0.3 1.19 0.76 2.01 19.25 0.23 0.19 1.43 0.92 2.41 2.09 0.27 0.19 1.65 1.06 2.80 0.00 0.31 0.19 1.86 1.19 3.19 0.00 0.35 0.19

0.4 1.17 0.75 1.99 23.22 0.22 0.19 1.40 0.89 2.39 4.60 0.26 0.19 1.61 1.03 2.78 0.00 0.30 0.19 1.81 1.15 3.16 0.00 0.34 0.19

0.5 1.16 0.73 1.98 26.75 0.22 0.19 1.37 0.87 2.37 4.60 0.26 0.19 1.61 1.03 2.78 0.00 0.30 0.19 1.75 1.11 3.12 0.00 0.33 0.19

0.6 1.14 0.72 1.97 30.96 0.22 0.19 1.34 0.85 2.35 5.23 0.25 0.19 1.52 0.96 2.72 0.84 0.30 0.19 1.69 1.07 3.09 0.00 0.33 0.20

0.7 1.12 0.71 1.96 32.43 0.21 0.19 1.30 0.83 2.33 10.25 0.25 0.19 1.48 0.90 2.68 4.18 0.29 0.20 1.64 0.93 3.05 0.84 0.33 0.20

Mean Minimum Maximum %Fu/FRk < 1 Standard D Variation

Table 1. Statistics obtained for the patch loading resistance of unstiffened girders, n = 478

αF0

λ–F0

0.25 0.5 0.75 1.0

0.2 1.19 0.75 1.68 6.78 0.17 0.14 1.43 0.91 2.04 0.42 0.20 0.14 1.65 1.05 2.38 0.00 0.23 0.14 1.85 1.19 2.48 0.00 0.24 0.14

0.3 1.17 0.74 1.66 10.59 0.17 0.14 1.40 0.89 2.01 0.85 0.20 0.14 1.60 1.02 2.33 0.00 0.23 0.14 1.79 1.15 2.50 0.00 0.25 0.15

0.4 1.15 0.72 1.64 13.14 0.16 0.14 1.36 0.87 1.92 1.69 0.19 0.14 1.55 0.99 2.27 0.42 0.23 0.15 1.72 1.11 2.47 0.00 0.26 0.16

0.5 1.12 0.71 1.62 19.07 0.16 0.14 1.32 0.84 1.94 2.54 0.19 0.15 1.49 0.96 2.22 0.85 0.23 0.16 1.66 0.99 2.50 0.42 0.27 0.17

0.6 1.10 0.70 1.60 25.42 0.16 0.14 1.28 0.82 1.90 3.81 0.19 0.15 1.44 0.93 2.17 1.69 0.23 0.16 1.59 0.95 2.43 0.85 0.27 0.17

0.7 1.08 0.69 1.58 36.44 0.16 0.14 1.24 0.80 1.86 5.80 0.19 0.14 1.40 0.90 2.12 2.68 0.23 0.14 1.54 0.96 2.36 1.79 0.26 0.13

Mean Minimum Maximum %Fu/FRk < 1 Standard D Variation

Table 2. Statistics obtained for the patch loading resistance of longitudinally stiffened girders, n = 770




This recalibration is carried out by assuming target valuesof the partial safety factor which are consistent with thevalues presently included within the core of the Eurocodesconcerning the strength and stability of steel structures.

References

[1] Lagerqvist, O., Johansson, B.: Resistance of I-girders to con-centrated loads. Journal of Constructional Steel Research 39(1996), No. 2, pp. 87–119.

[2] Graciano, C.: Patch Loading: Resistance of longitudinallystiffened steel girder webs. Doctoral Thesis, No. 2002:18,Lulea University of Technology, Sweden, 2002.

[3] EN 1993-1-5. Eurocode 3: Design of steel structures – Part 1-5:Plated Structural Elements, 2007.

[4] Davaine, L.: Formulations de la résistance au lancementd’une âme métallique de pont raide longitudinalement. Doc-toral Thesis, No. D05-05, INSA de Rennes, France, 2005 (inFrench).

[5] Gozzi, J.: Patch loading resistance of plated girders. Doc-toral Thesis, No. 2007:30, Lulea University of Technology,Sweden, 2007.

[6] Clarin, M.: Plate Buckling Resistance. Patch Loading ofLongitudinally Stiffened Webs and Local Buckling. DoctoralThesis, No. 2007:31, Lulea University of Technology, Sweden,2007.

[7] Chacón, R.: Resistance of Transversally Stiffened HybridSteel Plate Girders to Concentrated Loads. Doctoral Thesis,Universitat Politècnica de Catalunya, Barcelona Spain, 2008.

[8] Chacón, R., Mirambell, E., Real, E.: Hybrid steel plate girders subjected to patch loading. Part 2: Design proposal.Journal of Constructional Steel Research, 66 (2010), No. 5,pp. 709–715.

γ*M1 = 1.0 γ*

M1 = 1.1

λ–F0 αF0 λ–F0 αF0

0.5 1.0 0.5 0.75

Table 4. Values of the imperfection factor and the plateaulength to be used for achieving the desired level of safety

Step Feature Objective

Definition of the design mode FRk = cF · fyw · ly · tw To identify the basic variables

Experimental (or numerical) vs. predicted values

To compare the predicted values with the resistances obtained experimentally

and numerically

Estimation of the correction factor b r = b · ri,the · δTo create a probabilistic model of

resistance. “b” is the least square best fit of the slope, “δ” is the error term

Estimation of the variation of the error Vδ δii

the i

r

b r= exp,

,·

To determine the variation of the errors in the formulation which result from the

given formula

Normality test Kolmogorov-SmirnovTo test the normality of the distribution

of the errors

Definition of the variation of the fundamental variables Vxi

Mean Xi Vxi

To include a proper value (statistically determined) of the variations of the

fundamental variables

fy 1.14 · fy,nom 0.07

Thickness tnom 0.05

Width bnom 0.005

Definition of the variation of the numerical results VFEM

Numerical and experimental results are compared with each other to

determine the accuracy of the modelling of the steel girders.

This term is included for the sake of defining a certain level of random nature

in the calculation of resistance with deterministic numerical models

Definition of the final coefficient of variation Vr

V V V Vr Xkk

FEM2 2 2 21 1 1= +( ) +( )⎡

⎣⎢⎢

⎤

⎦⎥⎥

− +∏δ · To include all variations in a single formula

Definition of the characteristic resistance r b g X ek rt mk Q Q= − −∞· ( ) · ( · . · )0 5 2

kœ = 1.64 Q Vr= ( ) +ln 2 1

Definition of the new design value r b g X ed rt mk Q Qd= − −∞· ( ) · ( · . · ), 0 5 2

kd,œ = 3.04 Q Vr= ( ) +ln 2 1

Partial safety factor γM1 γ Mk

d

rr1 = To relate the characteristic to the design

resistances by means of a single coefficient

Corrected partial safety factor γ*M1 γ* ·M

k

d

n

k

rr

rr1 = To adapt the partial safety factor to better

statistical variations

Table 3. Procedure for determining the partial safety factor

re(exp)

rt (theoretical)




[9] Chacón, R., Bock, M., Real, E.: Longitudinally stiffened hy-brid steel plate girders subjected to patch loading. Journal ofConstructional Steel Research 67 (2011), No. 9, pp. 1310–1324.

[8] Müller, C.: Zum Nachweis ebener Tragwerke aus Stahl gegenseitliches Ausweichen. Doctoral Thesis, No. 47, RWTH Aachen,Shaker Verlag, 2003.

[11] Gabeler, L.: Statistical evaluation of patch loading resistancemodels for welded steel girders. Diploma Thesis, No. 2009-30X,Institute of Structural Design, University of Stuttgart, 2009.

[12] Viver, J.: Statistical evaluation of patch loading resistancemodels for longitudinally stiffened welded steel girders.Diploma Thesis, Technical University of Catalonia, 2011 (inSpanish).

[13] EN 1990: Eurocode – Basis of structural design, 2002.

Keywords: patch loading; EN 1993-1-5

Authors:Prof. Dr. Enrique Mirambell, Dr. Rolando Chacón: Construction EngineeringDepartment, Universitat Politècnica de Catalunya, C/Jordi Girona (CampusNord), 1-3, C1 Building, 08034 Barcelona, SpainProf. Dr. Ulrike Kuhlmann: Institute of Structural Design, University ofStuttgart, Pfaffenwaldring 7, 70569 Stuttgart, GermanyDr. Benjamin Braun: Federal Waterways Engineering & Research Institute,Kussmaulstr. 17, 76187 Karlsruhe, Germany

Day 1: Façades & structuresA wide range of projects was on offer,including active solar façades (Prof.Tina Wolf, Munich TU), challengingglass façades (Prof. Jens Schneider,Darmstadt TU), breathtaking, flexiblemembrane structures (Prof. Jan Cre-mers, Stuttgart HFT), composites in mo-bile structures (Prof. Göran Pohl, Uni-versity of Jena) and composites in com-petitive wall assemblies (Prof. JosefKurath, ZHAW, Winterthur). Withoutdoubt, composites have become indis-pensable, especially in smart façade systems.

Day 2: Natural forms & natural materialsClearly evident on the second day wasthe way the speakers employed a differ-ent approach to their treatment of themost diverse materials: the focus was onthe apparently negative properties, whichwere then used to advantage. Prof. JanKnippers (Knippers Helbig, Stuttgart/New York) presented shading louvresthat, owing to the low elastic modulus

Indispensable: fibre composites in architecture

4th International Symposium on Composites in ArchitectureWeimar, 5—6 December 2011“Structures & Skins”

It all began with the idea of bringing to-gether the world of fibre composites andthe world of construction in one place,a place steeped in history. This latestsymposium was the fourth time that topinternational architects and engineershad gathered in the charming ballroomof the “mon ami” youth and arts centrein Weimar to exchange experiences withfibre composite specialists. “Structures& Skins” was the topic this time.

As ever, the itinerary applied the con-cept of comparing innovations in build-ing with composites with innovations inother materials. Such an approach enablesa good overview of the future course ofthe construction industry and the contri-bution that fibre composites can make tothat course.

News

of the composite material, can adaptand deform to suit the position of thesun. This project is currently under con-struction (pavilion for Expo 2012, Yeosu,Korea; architects: SOMA Architecture,Vienna). Prof. Achim Menges (ICD, Stutt -gart TU) closed the day with fantasticcurves in timber – structures that derivefrom the same idea: exploiting the lowstiffness of the material.

As with the previous events, satisfied exhibitors and enthusiastic participantswere the hallmarks of the 4th Interna-tional Symposium on Composites in Ar-chitecture. In the meantime, the 100 orso delegates can be taken as an indica-tion that building with composites facesa bright future.

The next Composites in Architectureevent will take place on 3–4 December2012.

Dr.-Ing. Elke GenzelSKZ Das Kunststoff-Zentrumwww.skz.de/composites


16 © Ernst & Sohn Verlag für Architektur und technische Wissenschaften GmbH & Co. KG, Berlin · Steel Construction 5 (2012), No. 1

Articles

The corrugated steel plate has been used in many applications fora long time because of its favourable properties. Engineers haverealized a potential application in bridge structures, too, especiallyhybrid bridges. Among modern bridge erection methods, incremen-tal launching is one of the most competitive. When incrementallaunching is used to erect a bridge, the girder is subjected to acombined action of transverse force (F), shear force (V) and bend-ing moment (M), resulting in a complex stress field and a combinedloading situation. The current version of EN 1993-1-5, Annex D, in-cludes design methods for the bending and shear resistance ofcorrugated web girders, but no recommendations for the patchloading resistance calculation or consideration of different inter-actions (F+V; F+M). Therefore, this paper focuses on the experi-mental and numerical investigations into the structural behaviourof corrugated web girders under all the aforementioned actions.Design proposals are developed to determine the patch loadingresistance of corrugated web girders and two interaction equa-tions are proposed to take into consideration the reduction in re-sistance due to the combined loading situation.

1 Introduction

The corrugated steel plate is a structural element widelyused in many fields because of its many favourable proper-ties. This structural layout has spread in the field of bridges,too, especially for hybrid and composite bridges. In practice,during launching of a bridge structure, before the launch-ing device reaches the next pier, large shear, bending andtransverse forces can be introduced simultaneously at theprevious pier as shown in Fig. 1. In this case the corrugatedweb is loaded by a combination of large bending momentand shear force together with transverse forces in such a waythat their interaction should certainly be considered in thedesign. This paper focuses on the development of a designmethod for corrugated web girders under this combinedloading situation. The current version of EN 1993-1-5 [1]does not give recommendations for determining the patchloading resistance of corrugated web girders and there areno interaction equations for such girders. This paper sum-marizes the design proposals developed for the patchloading resistance and interaction equations which can beapplied in the design of corrugated web girders.

The previous experimental, analytical and numericalinvestigations have been analysed from the point of view ofbridge construction. These investigations were extendedto typical bridge structures and design methods developedwhich can be applied in the parameter range relevant tobridges. Patch loading tests were conducted at the BudapestUniversity of Technology & Economics, Hungary. The aimof the tests was to determine the patch loading resistanceof corrugated web girders with a typical corrugation pro-file of a real bridge girder under combined patch load andbending moment. Based on the experimental background,a finite element model was developed and verified. Cover-ing the wide range of parameters typical for bridges, acomprehensive parametric study was conducted to inves-tigate the patch loading resistance of corrugated web gird-ers. Based on the numerical model, the different interac-tion behaviours (F+V; F+M) were also studied. The differ-ent failure modes, load–deflection curves and ultimateloads were determined under each of the different loadingconditions. The structural behaviour under the combinedeffects was also analysed and the interaction of the differ-ent failure modes characterized. The intensity of the influ-ence of complex loading conditions was determined, eval-uated and design interaction equations developed to takeinto account the reduction in resistance due to the com-bined loading situation. The results reported in the cur-rent paper support the design of corrugated web girderbridges. The parameters applied are shown in Fig. 2.

Design of girders with trapezoidal corrugatedwebs under the interaction of patch loading, shearand bending

László DunaiBalázs Kövesdi*Ulrike KuhlmannBenjamin Braun

DOI: 10.1002/stco.201200003

* Corresponding author:e-mail [email protected]

Fig. 1. Internal forces during launchinga) F-V interaction: “pier reached” construction stageb) F-M interaction: “cantilever” construction stage

a)

b)

08_016-022_Koevesdi (2012-1-5-)_000-000_Steel Constr. (2sp).qxd 09.02.12 12:23 Seite 16


L. Dunai/B. Kövesdi/U. Kuhlmann/B. Braun · Design of girders with trapezoidal corrugated webs under the interaction of patch loading, shear and bending

2 Review of previous investigations2.1 Patch loading resistance of corrugated web girders

Experimental research on the patch loading resistance ofgirders with trapezoidal corrugated webs was started in1987 by Leiva-Aravena and Edlund [2]. Research in thisfield was continued by Kähönen in 1988 within the scopeof laboratory tests [3]. Based on the experimental investi-gations, a design model was developed to determine thepatch loading resistance of girders with corrugated webs.This design formula is based on the four plastic hinge fail-ure mechanism [1] developed by Rockey & Roberts for flatweb girders [4]. Kähönen modified this formula for corru-gated web girders [3] and the design method was cali-brated to the experimental results by several modificationfactors. Elgaaly and Seshadri carried out experiments in1997 to determine the patch loading resistance of corru-gated web girders [5]. A finite element model was devel-oped to analyse the patch loading failure mode and a nu-merical parametric study was conducted to examine theinfluence of the different parameters on the patch loadingresistance. A design method was proposed on the basis ofthe experimental and numerical investigations. Numericalinvestigations were also conducted by Luo and Edlund in1996 [6]. They proposed an empirical design formulabased on a simple design model for flat web girders.

2.2 Interaction of patch loading and shear force

Elgaaly presented the first experimental investigation avail-able on this topic for flat web girders in 1975 [7]. Zoete-meier conducted experiments on hot-rolled I-girders in1980 [8]. Experimental research activity was continued byOxfort and Weber on three welded test specimens in 1981[9]. Roberts and Shahabian carried out experiments on 24welded I-girders in 2000 and developed an interactionequation which is published in [10–12]. Kuhlmann andBraun researched this topic extensively in 2006 [13]. Basedon a large number of numerical calculations, a new inter-action curve was developed, which can be used over anextended parameter range. Elgaaly and Seshadri [5] first

investigated the shear and patch loading interaction of cor-rugated web girders in 1997. Five specimens were testedunder local transverse force. The main aim of their researchwas to determine the patch loading resistance of corrugatedweb girders. Their experiments led to the building of a nu-merical model. The pure patch loading resistance and theinteraction between bending, shear and patch loading wasstudied within the scope of a numerical research pro-gramme. Based on the numerical calculations, a design in-teraction curve was proposed in the form of Eq. (1):

(1)

where:VR shear buckling resistanceFR patch loading resistanceV applied shear forceF applied transverse force

2.3 Interaction of patch loading and bending moment

Research into this topic was started by Elgaaly in 1975 [7].Numerous tests were executed to determine the interac-tion behaviour and several interaction equations pro-posed. An overview of the previous investigations in thisfield for flat web girders can be found in [14]. Based on theprevious studies, the current EN 1993-1-5 [1] also includesan interaction equation for bending and transverse forcefor flat web girders. Elgaaly and Seshadri [5] first investi-gated the bending and transverse force interaction of thecorrugated web girders in 1997, in the aforementioned re-search programme. Based on a limited number of test re-sults and numerical calculations, an interaction equationwas proposed in the form of Eq. (2):

(2)

where:MR bending resistanceFR patch loading resistanceM applied bending momentF applied patch load

2.4 Summary of previous investigations and research aims

EN 1993-1-5, Annex D [1], deals with members with cor-rugated webs, but there is no standard design method fordetermining the patch loading resistance and there are nointeraction equations for corrugated web girders. The patchloading resistance and the interaction behaviour of corru-gated web girders was barely studied by researchers in thepast. Besides that, the results of previous investigations areof only limited use in bridge design because the majorityof them deal with shorter loading lengths and with higherglobal web ratios (hw/tw) than those typical for bridges. Somost of the tests executed are far from the typical parame-ter range of bridge structures.

Therefore, an extended analysis was executed withinthe scope of experimental and numerical investigationsand a design method is proposed to determine the patch

V FV

FFR R

− ⋅⎛

⎝⎜⎞

⎠⎟+

⎛

⎝⎜⎞

⎠⎟≤0 5 1 0

1 25 1 25. .

. .

FF

MMR R

⎛

⎝⎜⎞

⎠⎟+

⎛

⎝⎜⎞

⎠⎟≤

1 25 1 25

1 0. .

.

Fig. 2. Geometry of the trapezoidal corrugated girder




loading resistance of corrugated web girders. Furthermore,the structural behaviour of the corrugated web girders un-der the combined effect of bending, shear and patch load-ing is also investigated because there is only a very limitednumber of investigations available for corrugated webgirders. The applicability of the interaction curves (F+V;F+M) previously developed are studied and evaluated forbridge design, and new design equations proposed whichcan be used over a larger range of parameters.

3 Patch loading resistance of corrugated web girders

Patch loading tests were executed on 12 simply supportedgirders at the Department of Structural Engineering , Bu-dapest University of Technology & Economics, Hungary,in 2009. The aim of the tests was to determine the patchloading resistance of corrugated web girders with the typicalcorrugation profile of a real bridge girder. Ultimate loadswere determined, the structural behaviour and typical failuremodes analysed and described. The test arrangement andthe test specimen can be seen in Fig. 3. Details of the girderstested and the test results are published in [15] and [16].

A numerical model was developed and verified to ex-pand the applicability of the test results. The modellingwas based on a full shell model using four-node thin shellelements. Model verification was done by recalculatingmany experimental results. The ultimate loads were deter-mined by geometrical and material non-linear analysis us-ing imperfections. Based on the verified model, a compre-hensive parametric study was executed in order to analysethe patch loading resistance in the parameter range usedin bridges. The parameters analysed, which influence theload-carrying capacity, were varied between the followingvalues (for notation, see Fig. 2):– Web slenderness ratio: hw/tw = 200 to 500– Fold slenderness ratio: ai/tw = 13.3 to 116.7

(in all cases a1 = a2)– Loading length: ss/hw = 0.4 to 0.8 [2]– Corrugation angle: α = 15 to 65°

The details of the numerical model development, the modelverification and the detailed evaluation of the numericalsimulations are published in [15] and [16]. Numerical cal-culations were used for developing the design method which

may be applied to the practical design of corrugated webgirders. Based on the numerical database developed in[17], an enhanced design method was developed in a co-operation between the Institute of Structural Design, Uni-versity of Stuttgart, Germany (Braun and Kuhlmann), andthe Department of Structural Engineering, Budapest Uni-versity of Technology & Economics, Hungary (Kövesdi andDunai). The design method developed is based on the fourplastic hinge failure mechanism [3] according to the patchloading theory of Rockey and Roberts [4], which wasevolved for flat web girders. The design method was cali-brated for corrugated web girders by Kähönen [3] and fur-ther investigated and modified by Braun and Kuhlmannfor a larger parameter range [18]. Based on the own testresults and numerical calculations the design method isfurther developed by the authors and it can be applied ifthe failure mode is local buckling of the web. The geomet-ric criteria in Eq. (3) is to exclude the global buckling fail-ure of the whole web panel.

(3)

The first part of this equation refers to the resistance of theweb and the second part to the resistance of the flange.The web resistance can be calculated using Eq. (4):

(4)

The reduction factor χ should be calculated using Eqs. (5)–(9):

(5)

(6)

where:

(7), (8), (9)

fyw yield strength of webtw web thicknessai loaded fold length (if more folds are loaded, the max-

imum fold length)ss loading length

The new non-dimensional slenderness curve (λp–χ) was de-rived from Kähönen’s formula by Braun and Kuhlmann [18]and it can be applied to the patch loading resistance of cor-rugated web girders. The recommendation is illustrated inFig. 4. Curve ① is the proposal of EN 1993-1-5 [1] for patchloading of flat web girders (points show the results calcu-lated from tests), curve ② shows the proposal derived fromthe Kähönen formula and curve ③ is the recommended en-hanced non-dimensional slenderness curve for local buck-ling of corrugated web girders by Braun and Kuhlmann [18].Numerical results show that the corrugation angle has a sig-nificant influence on the patch loading resistance. Luo and

F F F for aht

tR R w R fl i

w

w

w= + ≥ +⎛

⎝⎜⎞

⎠⎟⋅, , .

26011 5

χλ λ

λ= − >1 9 0 798 1 2732

. . .p p

pif

λσ

σπ

υσ

pyw

crcr

w

i

f kE

ta

= =⋅

⋅ −⋅ ⋅

⎛

⎝⎜⎞

⎠⎟,

( ),

2

2

2

12 1kkσ = 1 11.

F t f ss kR w w yw, = ⋅ ⋅ ⋅ ⋅χ α

χ = 1 00. λ ≤ 1 273.if p

Fig. 3. Test setup




Edlund [6] developed an empirical design formula that con-siders the influence of the corrugation angle. Parametricstudies show that this curve does not follow the numericalresults in the case of relatively large corrugation angles(α > 40°). According to the numerical parametric study, thecharacter of the increasing patch loading resistance can bedescribed sufficiently well by referring the real length of thefolded web to the projected loading length. Therefore,Braun and Kuhlmann [18] propose improving the designmethod by introducing a factor kα, calculated from Eq. (10):

(10)

Fig. 5 shows the comparison of the two proposals and thenumerical calculations. The results of the numerical calcu-lations are presented by points, the proposal of Luo andEdlund is illustrated by curve ① and the proposal of Braunand Kuhlmann by curve ②. The results show that the pro-posal of Braun and Kuhlmann tracks the numerical resultswell and is on the safe side over the whole parameter rangeanalysed. It means that the applicability of the kα factordeveloped is proven based on the numerical calculations.

The flange contribution to the patch loading resistancemay be calculated using Eq. (11):

(11)

where:

(12)

ka aa aα =

++

1 2

1 4

F n M t fR fl pl f w yw, .= ⋅ ⋅ ⋅ ⋅ ⋅2 χ

Mb t

fpl ff f

yf. =⋅

⋅2

4

fyf yield strength of flangebf flange widthtf flange thicknessn plastic hinges considered in flange according to Table 1

Factor n depends on the tf /tw ratio. If the tf /tw ratio is < 4,the complete yield lines in the web and all four plastichinges in the flange can be evolved in the failure mecha-nism. If the flange is more dominant (tf/tw ≥ 4), the webfailure mechanism is activated earlier and none of the fourplastic hinges can develop, and so the patch loading resis-tance is smaller. This decrease in resistance can be takeninto account in the number of plastic hinges considered inthe flange and n can be determined according to Table 1.

Fig. 4. Non-dimensional slenderness and reduction factorrelationship

Fig. 5. Influence of the corrugation angle Fig. 6. Comparison between the numerical and experimen-tal results and the design method

Table 1. Number of the considered plastic hinges in theflange

tf/tw n

tf/tw < 4 4

4 ≤ tf/tw ≤ 7 3

7 < tf/tw 2

The new design proposal is compared with the resultsof the numerical and experimental database in Fig. 6. Thevertical axis of the diagram shows the numerical and ex-perimental results, the horizontal axis the results calculatedbased on the proposed design method. The comparisonshows a good correlation between the design method andthe numerical as well as experimental results. The meanvalue of the ratio of the results to the design method is1.008 and the standard deviation is 0.06. The design methodintroduced in this section determines the mean value ofthe patch loading resistance of the corrugated web girders.In order to be able to use this method in the design ofstructures, a partial factor is derived in [19] to determinethe design resistance level. Based on the statistical evalua-tion of the resistance model, a reduction factor of 1.35 canbe applied to reach the design resistance level. The partialsafety factor includes all the material and design methoduncertainties and already consists the γM1 = 1.1 what isrecommended in the EN1993-2 [20] for bridges.

The design method developed is valid for corrugationangles 15° ≤ α ≤ 65°, for loading lengths in the range ss/hw =0.4 and 0.8 [4] if inclined and parallel folds have the same




length (α1 = α2). This design formula can be used betweenweb ratios hw/tw = 200 to 500 and fold ratios α1/tw = 15 to100, if the failure mode is local buckling of one or moreweb folds. This parameter range is relevant for bridges.

4 Interaction between patch loading and shear force

During launching of a bridge structure, large shear andtransverse force can be introduced at the same cross-section.In this case, the corrugated web is loaded by large shear andtransverse forces, an interaction that should be consideredin the design. Numerical calculations have been carriedout to study the interaction behaviour of corrugated webgirders. The numerical model developed, the supports andthe loading conditions are introduced in Fig. 7. Details ofthe numerical model development and the model verifica-tion can be found in [20]. Different failure modes underpure patch loading, pure shear force and the combinedloading situations have been analysed in order to obtainthe structural behaviour of the shear and patch loading in-teraction. A detailed evaluation and comparison of the cal-culation results are published in [20].

The interaction behaviour of the corrugated web girderswas analysed in a large parameter range within the scopeof the numerical parametric study. The parameters investi-gated were as follows (for notation, see Fig. 2):– hw/tw = 100, 125, 150, 200 and 250– α1/tw = 15, 20, 25, 30 and 35– α = 20°, 30°, 40° and 60°– ss/hw = 0.2, 0.4, 0.5, 0.6 and 0.8

The ratio of F to V was changed for all the girder geometriesanalysed and calculations were executed with different F/Vratios in order to analyse the interaction behaviour. Theresults obtained, together with Elgaaly’s numerical calcula-tions, are presented in Fig. 8. The vertical axis of the dia-gram in Fig. 8 show the results as a function of (V-0.5F)/VR,num. The reason of it is that the shear stresses due to“pure patch load” are already included in the patch load-ing resistance model and a reduction of the load carryingcapacity is caused only by the additional shear stressescoming from “pure shear force”. The results show that the

interaction curve of Kuhlmann and Braun [13], whichwas developed for flat web girders, overestimates the inter-action behaviour for corrugated web girders. The recom-mended interaction curve of Elgaaly and Seshadri [5]gives a good approximation in the extended parameterrange as well. Of all the numerical calculations, only threeresults fall below this interaction curve and all of them areat the extremes of the parameter range analysed (longloading length ss, large α and large a1 at the same time). Inorder to take these special cases into account, a lowerlimit interaction curve is proposed in the form of Eq. (13). Itis recommended to use this interaction curve in the designof corrugated web girders:

(13)

Due to the relatively large scatter of the numerical results,parameters were studied separately to ascertain their ef-fects on the interaction criterion. The effect of the differ-ent geometric parameters are determined and evaluated in[19]. The analysis shows that the interaction behaviour ofcorrugated web girders depends on the FR,w/FR,fl ratio(web-to-flange patch loading resistance). This ratio can becalculated using Eqs. (4) and (11). Based on this assump-tion, a parameter-dependent interaction curve was devel-oped as an alternative, which takes into account the effectof the most influential parameters and therefore can leadto a more economic design. But this parameter-dependentinteraction curve does not consider the effect of the si-multaneous bending moment, which can reduce the safetyof the patch loading and shear interaction curve. There-fore, the application of this interaction curve is proposedonly in cases of small bending moments. The general use ofthis interaction curve needs more background investiga-tions with combined bending moments.

5 Interaction between patch loading and bending moment

During the launching of a bridge structure, large bendingmoments and transverse forces act in the same cross-sec-tion above the pier. The bending moment is mainly resistedby the flanges, and the shear and transverse forces are car-ried mainly by the web. But the flange loaded by the trans-verse force can have an important role in the patch loadingresistance as well, especially in the case of bridges where

V FV

FFR R

− ⋅⎛

⎝⎜⎞

⎠⎟+

⎛

⎝⎜⎞

⎠⎟≤0 5 1 0

1 2 1 2. .

. .

Fig. 7. Numerical model for investigating shear and patchloading interaction

Fig. 8. Interaction curve for shear and patch loading




the flanges are relatively thick. Therefore, the interactionof bending moments and transverse forces should also beconsidered in the design. Numerical calculations wereconducted to analyse the structural behaviour of the cor-rugated web girders under the combined loading. The typ-ical failure modes, load–deflection curves and load-carry-ing capacities were studied under transverse force, bend-ing moment and the combined loading condition. Thecomparison of the load–deflection curves calculated andthe load-carrying capacities in the interaction field arepresented in Figs. 9 and 10. The typical failure modes areweb crippling [5]and bending. Under the combined loadingsituation, the interaction of these two failure modes is typ-ical. A uniform transition between them can be observed inthe interaction zone, and no strict limit can be drawn. Theuniform transition can be also observed in the load–de-flection diagram in Fig. 10. Points (1) to (3) in Fig. 10 arefailures where patch loading dominates, and for points (4)to (7) it is bending that is dominant.

The interaction behaviour was analysed over a largeparameter range within the scope of a numerical paramet-ric study. The parameters investigated were the same asdescribed in Section 4. The results of the numerical andexperimental investigations are presented in Fig. 11. Thediagram shows that the combined bending moment andtransverse force hardly decreases the load-carrying capac-ity. The reason is the well-known “accordion effect” ofcorrugated web girders. Based on this effect, the bending

moment is mainly resisted by the flanges and the web car-ries the transverse force, and so the reduction in resis-tance is smaller than that for flat web girders. Besides that,the results show that the interaction curve according toEN 1993-1-5 [1] for flat web girders is on the unsafe side ifa large transverse force is acting on the girder with a smallbending moment. If both effects are relatively large, theEurocode curve is well on the safe side. These statementsare only valid if the reference values for the bending mo-ment resistance (MR_FEM) and the patch loading resistance(FR_FEM) are calculated by finite element analysis, and notby design formulas. The results show that the interactionproposal of Elgaaly and Seshadri [5] is conservative in theparameter range analysed; therefore, a new interactionequation (Eq. (14)) is proposed for corrugated web girders,which applies to the aforementioned parameter range:

(14)

Fig. 12 evaluates the results based on the bending momentresistance according to EN 1993-1-5 [1] and the patch load-ing resistance model proposed in this paper supplementedby the partial factor (1.35) determined in [19] for corrugatedweb girders. The comparison in Fig. 12 shows how themodel scatter of the aforementioned design formulas in-fluences the interaction. However, the statistical quality ofthe design formulas is not on the same level. The patchloading resistance model was developed on the basis of asmall experimental and numerical database, which results ina large partial factor. Therefore, the distance between thedata and the interaction curves is large.

6 Conclusions

Despite the increasing use of girders with trapezoidal cor-rugated webs in bridges, there are no design rules for de-termining the patch loading resistance nor its interactionwith bending moment and shear force. This paper outlinesdesign rule proposals that fit into the verification formatof EN 1993-1-5, Annex D [1].

The review of previous research work in this field hasrevealed not only a lack of data but also its focus on build-ing structures such that the parameters studied must be ex-panded to apply to bridge structures. The key achievementis the development of a patch loading resistance model in-

FF

MMR R

⎛

⎝⎜⎞

⎠⎟+

⎛

⎝⎜⎞

⎠⎟≤

1 20 8

1 0.

.

Fig. 9. Comparison of load–deflection curves

Fig. 10. Interaction behaviour

Fig. 11. Interaction curves, numerical and experimental results – based on resistances by FEM




cluding the statistical assessment of the partial factor so thata ready-to-use design rule can be proposed. Based on thenumerical and experimental database, interaction equa-tions for patch loading and bending moment (F-M) as wellas patch loading and shear force (F-V) are derived and thenewly developed patch loading resistance model success-fully evaluated and supplemented. Although the interac-tion between F-M and F-V is addressed in separate equa-tions at the moment, it is perceived that this is useful tothe designer because the internal force distribution duringbridge launching usually leads to patch loading either withdominant bending moment or dominant shear force. Nev-ertheless, future work will have to focus on merging bothinteraction equations in order to achieve a consistent def-inition of the F-M-V interaction surface.

Acknowledgements

The financial support provided by the German Academic Ex-change Service DAAD, by the Hungarian National ScientificResearch Foundation OTKA-T049305 project and by theNew Hungary Development Plan (project ID: TÁMOP-4.2.2.B-10/1-2010-0009) is gratefully acknowledged.

References

[1] EN 1993-1-5. Eurocode 3: Design of steel structures – Part1-5: Plated Structural Elements, 2007.

[2] Leiva-Aravena, L., Edlund, B.: Buckling of trapezoidally cor-rugated webs. Proceedings of the ECCS Colloquium on Sta-bility of Plates & Shells, Ghent, Belgium, 1987.

[3] Kähönen, A.: Zur Einleitung von Einzellasten in I-Träger mittrapezförmig profilierten Stegen. Stahlbau 57 (1988), No. 8,pp. 250–252.

[4] Roberts, T. M., Rockey, K. C.: A mechanism solution forpredicting the collapse loads of slender plate girders whensubjected to in-plate patch loading. Proceedings of the Insti-tution of Civil Engineers 67 (1979), No. 2, pp. 155–175.

[5] Elgaaly, M., Seshadri, A.: Girders with corrugated webs un-der partial compressive edge loading. Journal of StructuralEngineering 123 (1997), No. 6, pp. 783–791.

[6] Luo, R., Edlund, B.: Ultimate strength of girders with trape-zoidally corrugated webs under patch loading. Thin-WalledStructures 123 (1996), No. 6, pp. 135–156.

[7] Elgaaly, M.: Failure of thin-walled members under patchloading and shear. Proceedings of the 3rd International Con-ference on Cold-Formed Steel Structures, University of Mis-souri-Rolla, Montana, USA, 1975, pp. 357–381.

[8] Zoetemeijer, P.: The influence of normal-, bending- and shearstresses on the ultimate compression force exerted laterally toEuropean rolled sections. Report 6-80-5, Department of CivilEngineering, Delft University of Technology, Delft, 1980.

[9] Oxfort, J., Weber, N.: Versuche zum Stegblechbeulen bei Ein -zellasteinleitung und Biegung oder Querkraft. Report 7/1981,Institut für Stahlbau und Holzbau, Universität Stuttgart, 1981.

[10] Roberts, T. M., Shahabian, F.: Design procedures for com-bined shear and patch loading. Proceedings of the Institutionof Civil Engineers, Structures & Buildings 140 (2000), No. 3,pp. 219–225.

[11] Shahabian, F., Roberts, T. M.: Combined shear-and-patchloading of plate girders. Journal of Structural Engineering 126(2000), No. 3, pp. 316–321.

[12] Roberts, T. M., Shahabian, F.: Ultimate resistance of slen-der web panels to combined bending shear and patch loading.Journal of Constructional Steel Research 57 (2001), No. 7,pp. 779–790.

[13] Kuhlmann, U., Braun, B.: Combined shear- and patchload ing: numerical studies and development of an interactionequation. COMBRI project, RFCS Contract RFS-CR-03018,Document COMBRI-Report-USTUTT-002, 2006.

[14] Kövesdi, B., Dunai, L., Kuhlmann, U.: Bending momentand transverse force interaction acting on girders with corru-gated webs. Proceedings of the 6th European Conference onSteel & Composite Structures, EUROSTEEL 2011, vol. A, 31 Aug–2 Sept 2011, Budapest, Hungary, pp. 759–764.

[15] Kövesdi, B., Braun, B., Kuhlmann, U., Dunai, L.: Patch load -ing resistance of girders with corrugated webs. Journal of Con-structional Steel Research 66 (2010), No. 12, pp. 1445– 1454.

[16] Kövesdi, B., Dunai, L.: Determination of the patch loadingresistance of girders with corrugated webs using nonlinear finite element analysis. Computers & Structures 89 (2011),No. 11-12, pp. 2010–2019.

[17] Kövesdi, B.: The patch loading resistance of corrugated steelwebs used in bridge building. Diploma Thesis, No. 2007-20X,cooperation between the Institute of Structural Design, Uni-versity of Stuttgart, Germany and the Department of Struc-tural Engineering, Budapest University of Technology andEconomics, Hungary, 2007.

[18] Novak, B., Reichert, F., Röhm, J., Kuhlmann, U., Günther,H.-P., Braun, B., Raichle, J., Sprinke, P., Enders, W.: Berech-nungs- und Konstruktionsgrundlagen für sandwichähnlicheVerbundträger mit Trapezstegen im Brückenbau. Final Re-port, Project P 645, Forschungsvereinigung Stahlanwendung(FOSTA), Düsseldorf, 2008.

[19] Kövesdi, B.: Patch loading resistance of girders with corru-gated webs. PhD Dissertation, cooperation between Depart-ment of Structural Engineering, Budapest University of Tech-nology & Economics, Hungary, and Institute of StructuralDesign, University of Stuttgart, Germany, 2010.

[20] Kövesdi, B, Kuhlmann, U., Dunai, L.: Combined shearand patch loading of girders with corrugated webs. PeriodicaPolytechnica Civil Engineering 54 (2010), No. 2, pp. 79–88.

Keywords: corrugated web; patch loading resistance; shear andpatch loading interaction; bending and patch loading interac-tion

Authors:Prof. Dr. László Dunai, Dr. Balázs Kövesdi: Department of Structural Engineering, Budapest University of Technology & Economics, Műegyetem rkp. 3–5., 1111 Budapest, HungaryProf. Dr. Ulrike Kuhlmann: Institute of Structural Design, University of Stuttgart, Pfaffenwaldring 7, 70569 Stuttgart, GermanyDr. Benjamin Braun: Federal Waterways Engineering & Research Institute, Kussmaulstr. 17, 76187 Karlsruhe, Germany

Fig. 12. Interaction curves, numerical and experimental results – based on resistances by design formulas



The aim of this paper is to study the design requirements for rigidintermediate transverse stiffeners in longitudinally stiffened plategirders. Firstly, two representative design codes, AASHTO andEN 1933-1-5, are discussed and compared. Then the results oftwo tests on 1.5 m deep plate girders and the results of the exten-sive numerical parametric study are presented. The analysis ofthe results shows that the axial force in the stiffener due to ten-sion field action is overestimated in the design codes and that allthe design requirements for rigid transverse stiffeners can becovered by defining a minimum required second moment of area.

1 Introduction

Transverse stiffeners in plate girders are usually designedas rigid stiffeners that do not deflect beyond certain limitsand even at the ultimate limit state provide a solid supportalong the line of connection between the web plate of theweb itself and the longitudinal stiffener.

The main consequence of this is that the plate panelbetween adjacent rigid transverse stiffeners may be analysedas an isolated panel with well-defined boundaries. To beon the safe side, the panel is assumed to be restrained fortransverse movements and free to rotate along the edges.Further, the buckling length of longitudinal stiffeners iswell defined and can be assumed to be equal to the spac-ing of the transverse stiffeners.

Transverse stiffeners at end and intermediate sup-ports are subjected to large axial forces from support reac-tions. For this reason, they are mostly designed as double-sided stiffeners and are not dealt with in this paper. Inter-mediate transverse stiffeners are installed to increase thestrength and stiffness of the web panel and to prevent dis-tortional effects on the plate girder cross-section. Normally,they are designed as single-sided open stiffeners with arange of different cross-sections. The most typical are flat,L- or T-stiffeners.

In most cases, intermediate transverse stiffeners donot carry large external forces. If they are subjected to largeexternal forces, they have to be designed in a similar wayto support stiffeners, taking account of any eccentricity ifthey are single-sided. Usually, intermediate transverse stiff-eners are mainly subjected to forces arising from the fol-lowing two sources:

– Deviation forces from longitudinal stresses in the webpanels adjacent to the stiffener, which develop due to theout-of-plane imperfect geometry of the stiffener. Thesedeviation forces induce out-of-plane bending momentsin the stiffener and are subjected to second-order ef-fects.

– Tension field action that develops in the post-bucklingstate in shear. It develops in the form of a diagonal ten-sion band in the web plate and induces an axial com-pression force in the stiffener (truss analogy).

Both actions can appear individually or simultaneously. Toresist these actions and to limit deformability, rigid trans-verse stiffeners should, in principle, be designed for strengthand stiffness criteria at the ultimate limit state.

The first author to address the design issues of trans-verse stiffeners was Timoshenko [1]. Although his approachwas limited to geometrically perfect plates and stiffeners,the traces of his work can still be detected in current de-sign rules for deviation forces as well as for shear bucklingproblems.

Since 1950, many authors have dealt with transversestiffeners, but mainly for the effects of the shear loading inthe web panel: Stein & Fralich [2], Basler et al. [3], Rockeyet al. [4, 5], Evans et al. [6], Höglund [7], etc. Two designrules were typically developed: an expression for calculat-ing the minimum required second moment of area of thetransverse stiffener and an expression for the axial force inthe stiffener resulting from the tension field action in theweb panel, or, as an alternative, an expression for the min-imum required cross-sectional area of the stiffener takingaccount of local buckling but usually ignoring overall flex-ural buckling of the stiffener. A representative design codeof this kind is the American bridge design specificationAASHTO [8, 9], which will be discussed in more detaillater.

Recently, Euronorm EN 1993-1-5 [10] introduced amore general approach to stiffener design. According tothis, standard rigid intermediate transverse stiffeners shouldbe explicitly designed for strength and stiffness criteria,both at the ultimate limit state. Formulas for calculatingdeviation forces and the axial force in the stiffener are alsoprovided.

The common drawback of all the design rules men-tioned is that the axial force in the stiffener induced bytension field action in the post-buckling state is overesti-

Intermediate transverse stiffeners in plate girders

Franc Sinur*Darko Beg

Articles

DOI: 10.1002/stco.201200004


09_023-032_Sinur-Beg (SC7)_000-000_Steel Constr. (2sp).qxd 09.02.12 12:25 Seite 23

mated to a large extent, as has been demonstrated by sev-eral authors [11–15] and will be demonstrated again laterin this paper. This will be done based on two tests per-formed on 1.5 m deep plate girders and on the results ofan extensive numerical parametric study. The main aim isto find out if it is possible to cover all relevant design is-sues regarding the rigid intermediate transverse stiffenersof plate girders by the stiffness approach alone – definingthe minimum required second moment of area of the stiff-ener.

2 Brief overview of AASHTO and EN 1993-1-5 design standards

The two representative design standards, AASHTO andEN 1993-1-5, are discussed below.

2.1 AASHTO

In AASHTO-1996 [8], transverse stiffeners have to fulfilthe following second moment of area and cross-sectionalarea requirements:

(1)

(2)

where:hw depth of web platea spacing of transverse stiffenersc ratio of shear buckling resistance to shear yield strengthFyw specified minimum yield strength of stiffenerFcrs elastic local buckling stress of stiffener

The minimum stiffness requirement is needed for the webto develop the shear buckling resistance with near zerolateral deflection along the transverse stiffeners. The trans-verse stiffener has to resist the vertical component of thetension field action, and therefore Eq. (2) has to be ful-filled. This expression is derived from the condition thatthe maximum stress in the stiffener due to tension field ac-

Iha

atstw

w=⎛

⎝⎜⎞

⎠⎟−

⎛

⎝⎜⎜

⎞

⎠⎟⎟

max . ; .2 5 2 0 52

3

A Bh t c VV

tF

Fst w wu

wyw

crs

≥ − −⎡

⎣⎢

⎤

⎦⎥0 15 1 18 2. ( )


tion does not exceed the elastic local buckling stress in thestiffener. If the local buckling stress in the stiffener is higherthan the yield stress of the material, critical stress is re-placed with yield stress. In Eq. (2), the influence of eccen-tricity for single-sided stiffeners is taken into account withfactor B, which results in 1.0 for double-sided stiffeners,1.8 for single-sided angle stiffeners and 2.4 for single-sidedflat stiffeners. The effective part of the web plate, 18t2w, isalso accounted for in Eq. (2). In the minimum area condi-tion, only the post-buckling part of the force acting on thestiffener due to the tension field action is taken into ac-count.

According to AASHTO-1996 [8], the transverse stiff-ener of the longitudinally unstiffened web plate is only de-signed for the influence of tension field action. For longi-tudinally stiffened girders, Eq. (2) can be used to determinethe transverse stiffener area with hw equal to the wholegirder depth if the tension field develops over the wholeweb and hw equal to the maximum subpanel depth whentension fields in each subpanel develop independently.Additionally, the transverse stiffener used in the longitudi-nally stiffened web panels must also satisfy:

(3)

where Isl denotes the second moment of area of the longi-tudinal stiffener including an effective width of the webequal to 18tw and calculated for the neutral axis of thecombined section bst and bsl as shown in Fig. 1.

In the latest edition of AASHTO-2007 [9], the inter-mediate transverse stiffener that supports web subjectedto tension field action and deviation forces is given by themaximum of Eq. (1) and:

(4)

where:Ist second moment of area of transverse stiffener taken

around the edge in contact with the webρt maximum of Fyw/Fcrs and 1Fyw specified minimum yield strength of web plateFcrs local buckling stress for stiffener, given by:

Ibb

ha

Istst

sl

wsl≥

⎛

⎝⎜⎞

⎠⎟⎛

⎝⎜⎞

⎠⎟3

Ih F

Estw t yw≥

⎛

⎝⎜

⎞

⎠⎟

4 1 31 5

40ρ .

.

F. Sinur/D. Beg · Intermediate transverse stiffeners in plate girders

Fig. 1. Notation of the stiffened panel



(5)

In Eq. (5), bst and tst denote the width and thickness of thestiffener respectively, and Fys is the specified minimumyield strength of the stiffener.

2.2 EN 1993-1-5

According to Eurocode EN 1993-1-5 [10], the stiffener hasto meet strength and stiffness criteria given as:– The maximum stress σmax. in the stiffener should not ex-

ceed the yield strength:

– The additional lateral deflection wmax should not exceed:

When a transversely stiffened plate is loaded with pureshear only, the transverse stiffener has to fulfil minimumstiffness criteria:

(6a)

(6b)

This is similar to the stiffness requirement given by Eq. (1)of AASHTO [8, 9]. Due to tension field action in the webplate and truss analogy, the stiffener is subjected to an ax-ial force defined as:

(7)

where VEd is a design shear force and λw–––

. is the slender-ness of the web panel adjacent to the stiffener. Double-sided stiffeners may be simply designed as columns sub-jected to concentric compression. In the case of a single-sided stiffener, the stiffener should be verified for the axialforce and for bending moments due to the eccentricity ofthe axial force. Local buckling of open stiffeners is closelylinked to the torsional buckling mode and in EN 1993-1-5local buckling is prevented with the following expression:

σcr ≥ θfy (8)

where σcr is the elastic critical stress for torsional bucklingof the stiffener and θ is a parameter linked to the stiffener’s

Ih t

afor a

hstw w

w

≥ = ≤1 5

23 3

2

.α

I h t for ahst w w

w

≥ = >0 75 23. α

F E

bt

Fcrs

ts

st

ys=⎛

⎝⎜⎞

⎠⎟

≤0 312

.

σγmax ≤fy

M1

whw

max ≤300

N V thf

st ten Edw

wy

M, = − 1

321λ γ

cross-sectional shape (2 for flat stiffeners, 6 for stiffenersthat possess warping torsional stiffness).

Longitudinal compression stresses in the web plate andin longitudinal stiffeners due to bending moments and ax-ial forces induce transverse deviation forces. The magni-tude of these deviation forces qdev is related to the stiffenerimperfection amplitude w0 and is subjected to second-ordereffects. Fig. 2 shows the corresponding calculation modeladopted in EC 1993-1-5.

In most practical design cases, the web plate is sub-jected to a combination of shear and normal longitudinalstresses. The stiffener is subjected to an axial load Nst,tendue to tension field action and to deviation forces qdev dueto normal stresses acting on an initially imperfect stiffener.For these load combinations, the stiffener should be de-signed taking into account second-order effects, and bothstrength and stiffness criteria should be met. EN 1993-1-5does not explicitly prescribe the design checks for this gen-eral case, which is critical, especially for single-sided stiff-eners, because the overestimated axial force due to tensionfield action acts eccentrically on the stiffener.

A relatively simple design check that covers strengthand stiffness criteria, including second-order effects, is givenin Beg et al. [16] or Johansson et al. [17]).

Table 1 summarizes the minimum dimensions bst/tstof a flat stiffener calculated according to AASHTO-1996[8], AASHTO-2007 [9] and EN 1993-1-5 [10] for a selectedplated girder with dimensions hw/tw = 2000/8 mm, α =a/hw = 1 and steel grades S235 and S355. The actual loadacting on the stiffener is important only when the stiffeneris designed according to EN 1993-1-5. In this particular ex-ample, the stiffener was designed for the most unfavourable


Fig. 2. Calculation model for deviation forces in EN 1993-1-5

bst/tst (cm4)AASHTO (1996)

Eq. (2)AASHTO (2007)

Eq. (4)EN 1993-1-5

σmax,w

EN 1993-1-5Eq. (6)

hw = 200 cm, tw = 0.8 cm, a = 200 cm, S235

175/19 195/24 195/25 110/14

hw = 200 cm, tw = 0.8 cm, a = 200 cm, S355

190/20 230/28 205/25 110/14

Table 1. Transverse stiffener required to fulfil conditions given in EN 1993-1-5 and AASHTO


case, taking into account a shear load equal to the shearresistance of the web plate and a bending moment equalto the bending resistance of the web plate.

By increasing the yield strength, larger stiffener dimen-sions are required in all design provisions except in the caseof minimum stiffness criteria (Eq. (6) given in EN 1993-1-5).The latest AASHTO-2007 results in much larger stiffenersthan those of AASHTO-1996.

Recent studies regarding the design of transverse stiff-eners have shown that the formulation for the axial forcedue to tension field action (truss analogy) produces veryconservative values. The influence of the stiffness of thetransverse stiffener on a transversely stiffened girder undershear was studied by Lee et al. [12, 13]. In their study theyproved that the axial force in the stiffener is much smallerthan that predicted by design specifications. The result oftheir work is a new model that describes the post-bucklingbehaviour, called “shear cell analogy”, and a new proposalfor the stiffness requirement for a transverse stiffener.

The test results of the transverse stiffeners of Basleret al. [3] and Evans et al. [6] were evaluated by Höglund [11]and compared with the rules in EN 1993-1-5. The formu-lation in EN 1993-1-5 for determining the axial force fromtension field action was found to be conservative.


An extended numerical investigation into transverselystiffened girders has also been carried out by Handy &Presta [14, 15]. A new proposal for the axial force actingon a transverse stiffener was proposed:

FEd = VEd – Vbw,Rd ≥ 0 (9)

where VEd is the design shear load and Vbw,Rd is the de-sign shear resistance of the web only. This simplificationwas only proposed for transversely stiffened girders with-out longitudinal stiffeners.

3 Experimental work

Two tests on intermediate transverse stiffener were per-formed. Fig. 3 shows the layout of the girder tested andthe load positions for tests S1 and S2. The dimensions ofthe girder cross-sections are listed in Table 2. The interme-diate transverse stiffeners were designed for the effects ofdeviation forces and tension field action. The deviationforces were calculated from the stress distribution due tothe pure bending resistance of the plate. Only half of theaxial force due to tension field action, calculated accord-ing to Eq. (7), was considered in the design. In both cases


Fig. 3. Layout of the girder tested and loading positions for tests S1 and S2

Web Top flange Bottom flange Longitudinal stiffener

Specimen hw [mm] tw [mm] a [mm] bf1 [mm] tf1 [mm] bf2 [mm] tf2 [mm] Hsl [mm] hsl [mm] bsl [mm] tsl [mm]

SO 1500 7 1500 320 22 320 22 – – 90 10

SC 1500 7 2250 320 22 320 22 160 80 80 5

UO 1800 6 1800 250 20 450 20 – – 100 10

UC 1800 6 2700 250 20 450 20 300 180 80 5

Table 2. Geometry of the steel plate girders tested



studied, the dimensions of the transverse stiffener were fi-nally set to bst/tst = 120/15 mm.

The out-of-plane displacements in the panel and inthe transverse stiffener investigated were measured at dis-crete points by displacement transducers. Strains in thetransverse stiffener and the web plate in the vicinity of thestiffener were measured in addition to displacements, asshown in Fig. 4. The strain gauges were positioned on eitherside of the stiffener and the web plate.

The initial geometrical imperfection of the transversestiffener was measured and in both cases bow imperfectionin the direction of the intermediate transverse stiffener wasdetected, with a maximum amplitude of 4.7 mm for stiffenerS1 and 3.7 mm for stiffener S2.

3.1 Test results

The load-deflection curves for both tests are plotted inFig. 5. In the first test (S1), the load was increased up to2572 kN. The test was stopped just before the maximumcapacity of the girder was reached. In the second test (S2),the test specimen was loaded up to the maximum resis-tance of 2659 kN – well into the softening range in termsof displacements.

The average strains over the thickness of the transversestiffener are plotted in Fig. 6. The resulting maximum mem-brane stress for stiffener S1 is obtained at the position ofstrain gauge L12 in the web plate with a value of –210 MPa,which is less than the measured yield stress of the material.In test S2, plastic strains were observed at point L11 only,with a maximum average strain of 0.46 %. This strain wasobtained after the peak load had already been attained.The strain at maximum load was 0.30 %. In both tests, themaximum average strains were measured in the cross-sec-tion near the longitudinal stiffener. This is due to the factthat this cross-section is directly in the area where the di-agonal tension field is anchored into the transverse andlongitudinal stiffener. In this cross-section (1-1, see Fig. 7)the strains are relatively high, whereas in the other twocross-sections (2-2 and 3-3) the strains are smaller and also


Fig. 4. Positions of strain gauges onstiffener and web

Fig. 5. Load-deflection curves for tests S1 and S2

more representative for determining the axial force in theintermediate transverse stiffener.

The axial forces were evaluated from the measuredstrains for each cross-section. The results are listed inTable 3. For comparison, 100 % (100 % TFA) and 50 %(50 % TFA) of the axial force due to tension field actionaccording to EN 1993-1-5 are given. As can be seen, themaximum compression is found in section 1-1 (see Fig. 7).One of the measured points (L11) was directly in this di-agonal tension field where the strains are extremely high.In the middle section (2-2), the axial force is much smaller,whereas section 3-3 contains the smallest value. It is rea-sonable to assume that section 2-2 is relevant for deter-mining the representative (average) value of the axial forcein the stiffener. This axial force represents only 56 % ofthe calculated axial force arising from tension field actionusing truss analogy.

4 Numerical simulations4.1 Model verification

The numerical model was developed in the multi-purposeprogram ABAQUS [18] and was verified against the test re-sults. The measured initial geometrical imperfections and

Nten [kN] Stiffener S1 Stiffener S2

SECTION 1-1 2-2 3-3 1-1 2-2 3-3

TEST – 329.1 – 290.0 – 223.4 – 653.9 – 280.7 – 160.4

100 % TFA – 514 – 504

50 % TFA – 257 – 252

Table 3. Axial force in transverse stiffener at maximum girder resistance, taking into account effective part of web 15εtw



non-linear material behaviour based on tensile tests wereconsidered. Verification of the numerical model was per-formed by comparing the initial stiffness, maximum capacityand load-deflection curve.

Fig. 8 compares the experimental and numerical re-sults by way of the load-deflection curves. The initial stiff-ness of numerical model S1 is slightly higher than the ex-perimental one. The transition from the elastic to the plas-


Fig. 7. Cross-sections in stiffener where axial forces wereevaluated

Fig. 8. Comparison of load-deflection curves for test S1

Fig. 6. Strain measurements in transverse stiffeners



tic zone is very similar, but the maximum capacities cannotbe compared because the test was stopped before the resis-tance was reached. However, comparing the load obtainedat the same vertical displacement, the difference betweennumerical and experimental values is small. As noted in theprevious case, the initial numerically obtained stiffness oftest S2 (see Fig. 9) is also slightly higher than the experi-mental one. The calculated resistance is 3.7 % lower com-pared with the experimental results.

4.2 Parametric study

Two sets of numerical analyses were performed. First of all,the influence of the stiffness of transverse stiffeners wasstudied on girders that were equivalent to the test girdersS1 and S2. The second series was performed on girdersloaded with a combination of high bending and high shearloads (see Fig. 10). All important parameters were variedin the analysis.

A bilinear material model with yield strength fy =355 MPa was used in the parametric studies. According toEN 1993-1-5, the equivalent geometric imperfection of thetransverse stiffener under consideration was taken as ahalf sine wave imperfection with amplitude w0 = hw/300.Each adjacent stiffener was straight (see Fig. 2). Both im-perfection directions were studied. It was found that theimperfection in the direction of the transverse stiffener isdecisive due to larger eccentricity of the tension field ac-tion.

4.2.1 Stiffeners S1 and S2

The size of the transverse stiffener stiffness was varied fornumerical simulations on test girders S1 and S2. Table 4summarizes the dimensions and stiffnesses of the stiffenersused. The stiffness of the stiffener was normalized with theminimum required stiffness (see Table 5) given by Eq. (6)and with the stiffness required to fulfil strength and stiff-ness requirements taking account of tension field action atthe ultimate shear resistance of the girder and deviationforces due to the corresponding bending moment in theadjacent panel. The simplified stiffener resistance modelfrom Beg et al. [16] was used for this purpose.

The results of the numerical simulations are presentedthrough:a) The evolution of out-of-plane displacements along the

transverse stiffener.


Fig. 10. Numerical model of stiffened plate girder under combination of high bending moment and shear force

Fig. 9. Comparison of load-deflection curves for test S2

TEST\STIFFENER I1 I2 I3 I4 I5 I6 I7 I8

S1- bst/tst [mm] 60/6 70/7 80/8 110/11 120/12 130/13 140/14 150/15

I/Ireq Eq. (6) 0.33 0.58 0.94 2.90 3.92 5.15 6.63 8.37

I/Ireq 100 % TFA 0.07 0.12 0.20 0.61 0.83 1.09 1.40 1.77

I/Ireq 50 % TFA 0.21 0.37 0.60 1.86 2.52 3.31 4.26 5.38

S2- bst/tst [mm] 60/6 70/7 80/8 100/10 120/12 140/14 150/15 158/16

I/Ireq Eq. (6) 0.23 0.40 0.65 1.44 2.72 4.60 5.81 6.99

I/Ireq 100 % TFA 0.07 0.13 0.21 0.46 0.86 1.46 1.85 2.22

I/Ireq 50 % TFA 0.24 0.42 0.67 1.49 2.80 4.74 5.98 7.19

Table 4. Stiffness variation of transverse stiffeners in tests S1 and S2

Ireq (cm4) EN 1993-1-5 Eq. (6) 100 % TFA 50 % TFA

S1 115.2 545.4 179.4

S2 165.9 522.4 161.2

Table 5. Stiffness of stiffener according to different require-ments


b) The amplitude of the maximum displacement and max-imum capacity for transverse stiffeners with differentstiffnesses.

Fig. 11 shows the out-of-plane displacements obtained atthe maximum force for cases S1 and S2. The shape of theout-of-plane displacement is directly linked with the stiff-ness of the stiffener. The “S” shape of the out-of-plane dis-placement of the transverse stiffeners was observed forstiffnesses I1 and I2. These stiffeners are not rigid so theydo not prevent shear buckling over the depth of the web.The lateral deflection over the height of the stiffener isgoverned by the buckling in the panel. The amplitudes forweak stiffeners were –10.5 and 6.2 mm for cases S1-I1 andS2-I1 respectively. By increasing the stiffness of the trans-verse stiffener, the shape was transformed from the “S” tothe “C” (I3) shape with much smaller amplitudes.

Fig. 12 shows the influence of the stiffness of trans-verse stiffener on the maximum resistance and maximumamplitude of the out-of-plane displacement. On the hori-zontal axis, the actual stiffness of the transverse stiffener isnormalized with stiffness Ireq calculated as:a) Minimum stiffener requirement according to Eq. (6).b) Required stiffness calculated to fulfil strength and stiff-

ness requirements under the influence of deviationforces and tension field action (EN 1993-1-5).

c) Required stiffness calculated to fulfil strength and stiff-ness requirements under the influence of deviation forcesand 50 % of tension field action (EN 1993-1-5).

An instantaneous decrease in the maximum out-of-planedisplacements and increase in the girder capacity of gird-ers S1 and S2 had already been found at small I/Ireq ra-tios. The increase in the girder capacity plotted on the ver-tical axis in Fig. 12 was very small. The difference betweenmaximum and minimum values was only 3.5 % for girderS1 and 2 % for girder S2. As can be seen from the diagram,the maximum displacement of the stiffener is already belowthe hw/300 limit for the small stiffener I2 (see Table 4). To


fulfil the displacement criterion (w ≤ hw/300) in case a),40 % of the required stiffness is needed for girder S1 and60 % for girder S2, in case b), 13 % for S2, and in case c),42 % for S2. In all cases the stiffness requirement was foundto govern. The stresses in the stiffener were always belowyield strength, except in the part of the web where the ten-sion field formed.

4.2.2 Stiffeners subjected to high bending moment and shear force

The parameters in the parametric analysis were: stiffness oftransverse stiffener, web slenderness, ratio of flange area toweb area, stiffness of longitudinal stiffener and panel aspectratio. Seven different girder cross-section geometries wereanalysed. The basic geometry of the girder (G1 in Table 6)is defined with the following parameters: hw/tw = 250,Af/Aw = 0.7, γ/γ* = 3.0 and α = a/hw = 1.0, where γ denotesthe relative stiffness of the longitudinal stiffener and γ* thelimit stiffness of the longitudinal stiffener which preventsoverall buckling of the web panel in pure shear, including


STIFFENERbst/tst [mm]

I1 I2 I3 I4 I5 I6 I7 I8

20/2 40/4 60/6 80/8 100/10 120/12 150/15 200/20

GIRDERG1 G2 G3 G4 G5 G6 G7 G8

hw/tw = 250 Af/Aw = 0.3 Af/Aw = 1.1 γ/γ* = 0.30 γ/γ* = 1.00 α = 0.5 hw/tw = 150 hw/tw = 350

Table 6. Parameters taken into account for girders loaded with high bending and shear load

Fig. 11. Out-of-plane displacement of transverse stiffeners S1 and S2

Fig. 12. Influence of stiffness of transverse stiffener on girderresistance and out-of-plane displacement of stiffener



deflection of this stiffener. In all other cases (girdersG2–G8), only one parameter was changed compared withthe basic girder (G1). Within each girder, the additional


Ireq (cm4) AASHTO Eq. (1) EN 1993-1-5 Eq. (6) 100 % TFA 50 % TFA

G1–G5 51.2 153.6 3617.4 887.1

G6 409.6 614.4 1696.0 697

G7 237.0 711.1 710.2 349.6

G8 18.7 56.0 2529.7 753.8

Table 7. Required stiffener stiffnesses considering different requirements

Fig. 13. Out-of-plane displacement along transverse stiffenerfor girder hw/tw = 250, α = 1, γ/γ* = 3.0, Af/Aw = 0.7

Fig. 14. The normalized force obtained at an out-of-plane displacement of hw/300 for different stiffener stiffnesses

parameter was the stiffness of the transverse stiffeners. Thegeometry of the transverse stiffeners in the parametric studyis given in Table 6.

Table 7 lists the second moments of area for all trans-verse stiffeners (I1–I8) calculated for different code require-ments – AASHTO (Eq. (1)), EN 1993-1-5 (Eq. (6)) – and inthe last two columns according to strength and stiffnesschecks assuming deviation forces and forces due to tensionfield action. The column “100 % TFA” denotes that fulltension field action was considered, and “50 % TFA” de-notes that only 50 % of tension field action was taken intoaccount in the design of the stiffener.

The out-of-plane deflections of transverse stiffenersloaded with a combination of shear force and bendingmoment are plotted in Fig. 13. The curves are plotted forgirder G1 and for different transverse stiffeners (see Table 6).The deflected shape depends on the stiffness of the stiff-ener. By increasing the stiffness, the deflection of the stiff-ener is transformed from the “S” shape to the “C” (I5)shape. The resistances of the girders obtained at a deflec-tion of hw/300 = 6.67 are plotted in Fig. 14. The resistance


was normalized with the maximum force obtained in allgirders with the same cross-section properties which wereanalysed, whereas the actual stiffness is normalized withthe required stiffness given by Eq. (6) and with the stiff-ness required to fulfil strength and stiffness conditions tak-ing into account 50 or 100 % of the tension field action.When the value 1 is reached on the y axis, the resistanceof the girder proves its maximum resistance and at thesame time the out-of plane deflection of the intermediatetransverse stiffener is below the limit value of hw/300. Atthis point, minimum stiffness can be read to fulfil stiffnesscriteria. As it can be seen, the EN 1993-1-5 stiffness re-quirement generally covers most of design cases becausethe value 1 on the y axis is reached before value 1 on the xaxis.

5 Conclusions

If the transverse stiffener of a plate girder is designed toaccommodate deviation forces from bending momentsand axial forces from tension field action according toEN 1993-1-5, this results in much bigger stiffener thanthat obtained from numerical simulations. This can be at-tributed to the overestimation of the axial forces in thestiffener due to tension field action. The actual force, mea-sured in our own tests, represents 56 % of the force calcu-lated according to Eq. (7). A similar effect was also ob-served in [12–15].

This overestimation of the axial force in the stiffeneris very important for the single-sided stiffeners mostly usedfor intermediate transverse stiffeners. It has been demon-strated in this paper that the required performance of rigidintermediate transverse stiffeners of longitudinally stiffenedplate girders may be obtained by fulfilling simple stiffnesscriteria given by Eq. (6) increased by a factor that is nolarger than 3 but can in most cases be equal to 1. This sim-plifies the design procedure for intermediate transversestiffeners while considering all relevant effects.

References

[1] Timoshenko, S.: Theory of Elastic Stability, McGraw-HillBook Company, New York/London, 1936.

[2] Stein, M., Fralich, R. W.: Critical shear stress of infinitelylong simply supported plate with transverse stiffeners. Jour-nal of Aeronautic Science, vol. 17, 1950, p. 38.

[3] Basler, K. et al.: Web buckling tests on welded plate girders.Welding research council, Bulletin No. 64, New York, 1960.

[4] Rockey, K. C.: An Ultimate Load Method for the Design ofPlate Girders. In: Proceedings of Colloquium on design ofplate and box girders for ultimate strength. London, 25–26Mar 1971, IABSE, pp. 253–268.


[5] Rockey, K. C., Evans, H. R., Porter, D. M.: Ultimate LoadCapacity of Stiffened Webs Subjected to Shear and Bending.In: Proc. Conf. Steel Box Girders. London, 1973.

[6] Evans, H. R., Tang, K. H.: A report on five tests carried outon a large-scale transversely stiffened plate girder – TRV3.University College, Cardiff, 1981.

[7] Höglund, T.: Shear buckling resistance of steel and aluminiumplate girders. Thin-Walled Structures, vol. 29, 1997, pp. 13–30.

[8] AASHTO: Standard specifications for highway bridges.Washington, D.C., American Association of State Highway &Transportation Officials, 1996.

[9] AASHTO: Standard specifications for highway bridges.Washington, D.C., American Association of State Highway &Transportation Officials, 2007.

[10] EN 1993-1-5. Eurocode 3: Design of steel structures –Part 1-5: Plated structural elements. Brussels: EuropeanCommittee for Standardization, 2006.

[11] Private communications with T. Höglund, 2005.[12] Lee, S. C., Yoo, C. H., Yoon, D. Y.: Behaviour of Interme-

diate Transverse Stiffeners Attached on Web Panels. Journalof Structural Engineering, vol. 128, 3, 2002, pp. 337–345.

[13] Lee, S. C., Yoo, C. H., Yoon, D. Y.: New design rule for In-termediate Transverse Stiffeners Attached on Web Panels.Journal of Structural Engineering, vol. 129, 12, 2003,pp. 1607–1614.

[14] Hendy, C. R., Presta, F.: Transverse web stiffeners and shearmoment interaction for steel plate girder bridges. In: Inter -national Conference on Steel Bridges. Guimarães, Portugal,4–6 Jun 2008. ECCS, p. 8.

[15] Presta, F.: Post-buckling behaviour of transversely stiffenedplate girders. Doctoral Thesis. Cosenza, Universita degli studidella Calabria, 2007, p. 164.

[16] Beg, D., Kuhlmann, U., Davaine, L., Braun, B.: Eurocode 3:Design of Steel Structures, Part 1-5: Design of Plated Struc-tures. 1st. ed, ed. ECCS, vol. 1. Berlin: Ernst & Sohn, 2010,p. 272.

[17] Johansson, B., Maquoi, R., Sedlacek, G., Müller, C., Beg, D.:Commentary and worked examples to EN 1993-1-5 “PlatedStructural Elements”. In: JRC Scientific and Tehnical Reports,2007.

[18] SIMULIA: Abaqus Online Documentation: version 6.7.EF1. In: Deassault Systemes, 2008.

Keywords: plate girders; transverse stiffeners; tension field; de-viation forces; experimental tests; numerical simulations

Authors:Dr. Franc Sinur: University of Ljubljana, Slovenia, Faculty of civil and geodetic engineering,chair of Metal Structures, Sloveniae-mail: [email protected]

Prof. Darko Beg: University of Ljubljana, Slovenia, Faculty of civil and geodetic engineering,chair of Metal Structures, Sloveniae-mail: [email protected]




European standard EN 1993-1-5 provides the effective width methodand the reduced stress method as calculation methods for platebuckling assessment. This article focuses on the reduced stressmethod, for which the global slenderness concept was introducedinto EN 1993-1-5. The concept offers advantages associated withtoday’s numerical procedures which allow the elastic critical loadfactor of the full stress field to be determined in a single step, thustaking interaction into account in a very elaborate way. However,this article shows that the interaction verification in its pure for-mat based on the von Mises yield criterion is not able to repre-sent the actual mechanical behaviour of biaxially compressedplates. A simple modification is proposed which leads to appro-priate and plausible results.

1 Introduction

Slender plates and plated structures subjected to compres-sive in-plane loadings are intrinsically sensitive to buck-ling. As buckling is a complex process that involves non-linear stress distributions, load-shedding effects betweencross-sectional parts and the influence of imperfections aswell as boundary conditions, even today, designers usuallyfavour easy-to-handle calculation methods over non-linearnumerical analyses. European standard EN 1993-1-5 [3]offers two different calculation methods for such a platebuckling assessment: the effective width method and thereduced stress method. The former is highly efficient forstandard geometries because it accounts for load-sheddingbetween cross-sectional parts. As a requirement, it subdi-vides the stress field into stress components for which in-dividual plate buckling checks are required (EN 1993-1-5,chapters 4 to 6) and which are finally combined in inter-action rules (chapter 7). However, the effective widthmethod is not applicable for non-uniform geometries andcertain types of loading. In contrast, the reduced stressmethod applies to almost any geometries and loadingsdue to the generic concept that takes into account the fullstress field and its interaction (EN 1993-1-5, chapter 10).This article focuses on the reduced stress method andsolves shortcomings for biaxially compressed plates inparticular.

2 Design methodology2.1 General

Although the elastic critical buckling load has been thestarting point for an understanding of plate buckling be-haviour, linear buckling theory alone is not an appropriatepredictor of the ultimate buckling strength. The asset oflinear buckling theory is, however, that the elastic criticalbuckling load can be explicitly mathematically defined, as-suming perfect geometry and linear material behaviour. Thismakes it, even today, well suited as an input parameter forthe common approach to determine the ultimate bucklingstrength via a reduction in plastic strength:

(1)

where:Rpl plastic strengthRcr elastic critical buckling strengthRult ultimate buckling strengthλ– slendernessρ(λ–) reduction factor, function of the slenderness

The reduction factor ρ (also denoted χ or κ) encompassesthe influence of structural and material imperfections. More-over, it reflects the considerable post-critical strength reserveand the detrimental column-like behaviour that plates canpossess. In EN 1993-1-5 [3], the effective width and reducedstress methods make use of the design methodology in dif-ferent ways. Since from the European viewpoint the re-duced stress method is often perceived as a German spe-ciality, the following section tries to give an overview on theprocedure and background of the concept in EN 1993-1-5,chapter 10 as well as of DIN 18800-3.

2.2 Reduced stress method2.2.1 EN 1993-1-5 concept

EN 1993-1-5, chapter 10 [3], provides the framework for thereduced stress method in sections 10(1), 10(3), 10(4) and10(5b). Section 10(7) may also be relevant for stiffener de-sign and detailing. The basic idea behind the global slen-derness concept in EN 1993-1-5 follows former Germanguideline DASt-Ri 012 [1] dating from 1980. It differs from

λ ρ λ= → = ( )R

RR Rpl

crult pl·

Reduced stress design of plates under biaxial compression

Benjamin Braun*Ulrike Kuhlmann

Articles

DOI: 10.1002/stco.201200005


10_033-040_Braun (2012-1-4)_000-000_Steel Constr. (2sp).qxd 09.02.12 12:32 Seite 33

German standard DIN 18800-3 [2] which is currently inuse, see section 2.2.2.

Section 10(1) explains the application range in brief.At present, load-shedding between cross-sectional parts inplated structures is not accounted for. Thus, the weakestpart of the cross-section usually governs the resistance ofthe whole cross-section, which may lead to significant dif-ferences when results are compared with the effectivewidth method. However, it was shown in [11] that the re-duced stress method could be expanded to overcome this.

Section 10(3) introduces the global slenderness λ–p:

(2)

where:

αult,k ; characteristic resistance load factor

αcr ; elastic critical load factor

fy yield strengthσeq,cr equivalent elastic critical buckling stressσeq,Ed design equivalent stress

Expressed in plain terms, the global slenderness relates theyield strength fy to the equivalent elastic critical bucklingstress σeq,cr. The reduced stress method is used most effi-ciently when σeq,cr can be determined with numerical pro-cedures in one step. Since the output of a linear bifurcationanalysis is commonly a load factor instead of a stress value,the elastic critical load factor αcr has been introduced forconvenience. Section 10(6) describes a manual calculationprocedure for αcr which is, however, quite tedious. By re-lating the elastic critical load to the design equivalent stressσeq,Ed, the same has to be done for the characteristic resis-tance. However, both quotients can be reduced in the end,see Eq. (2).

Determining the global slenderness for unstiffenedplates usually means performing a single calculation for thefull stress field. For stiffened plates, at least two separatecalculations are required which take into account eitherlocal buckling of subpanels or global buckling of the stiff-ened panel.

In the next step, the global concept fades into a semi-global one because the determination of the reduction fac-tors and the evaluation of column-like buckling are attributedto the corresponding stress component, see Eq. (3). The de-signer has to choose the appropriate reduction function de-pending on the boundary conditions that represent the sys-tem. In EN 1993-1-5, this means choosing from section 4.4(2)and/or Annex B for plate buckling, 4.5.3(5) for column buck-ling and 5.3(1) and/or Annex B for shear buckling:

(3)

where:ρx reduction factor for longitudinal stresses, interpolated

value ρx,c where column-like buckling is relevant

λαα σpult k

cr

y

eq cr

f= =,

,

=fy

eq Edσ ,

=σσ

eq cr

eq Ed

,

,

ρ λ ρ λ χ λx p z p w p( ) ( ) ( ); ;


ρz reduction factor for transverse stresses, interpolatedvalue ρz,c where column-like buckling is relevant

χw reduction factor for shear stresses

The interpolation between plate-like and column-like buck-ling follows section 4.5.4(1) of EN 1993-1-5. It is based onthe ratio between the elastic critical plate buckling stressσcr,p and the elastic critical column buckling stress σcr.c.The determination of σcr,p and σcr.c is carried out for eachrelevant stress component separately assuming identicalboundary conditions.

The general verification format according to Eq. (4) isgiven in section 10(5b). The interaction between stress com-ponents is based on a variation of the von Mises yield cri-terion:

(4)

whereσEd; τEd design load valuesσx,Rd = ρx · fyd; design resistance valueσz,Rd = ρz · fyd; design resistance valueτRd = χw · fyd; design shear resistance valuefyd design yield strength

Sections 10(2) and 10(5a) offer an alternative verificationformat according to Eq. (5). Both sections correspond witheach other and it is presumed that the minimum value ofall reduction factors is taken. Without doubt, it may beconvenient in the preliminary dimensioning process but itis usually too conservative in the final design of plates whendifferent loadings are present:

(5)

where:σeq,Ed design equivalent stressρmin = min (ρx;ρz;χw); minimum value of all reduction fac-

torsfyd design yield strength

2.2.2 DIN 18800-3 concept

The reduced stress method is incorporated in a different wayin German standard DIN 18800-3 [2]. In the following, DINnotation has been slightly adjusted to compare with ENnotation and only the most prominent differences are out-lined. Instead of the semi-global concept of EN 1993-1-5,DIN 18800-3 focuses on stress components from the be-ginning. Thus, a separate slenderness for each relevantstress component is determined:

(6)

where:fy characteristic yield strengthσx,cr; σz,cr; τcr elastic critical buckling stress

σσ

σσ

σσ

x Ed

x Rd

z Ed

z Rd

x Ed

x

,

,

,

,

,⎛

⎝⎜

⎞

⎠⎟ +

⎛

⎝⎜

⎞

⎠⎟ −

2 2

,,

,

,

· ·Rd

z Ed

z Rd

Ed

Rd

⎛

⎝⎜

⎞

⎠⎟

⎛

⎝⎜

⎞

⎠⎟ +

⎛

⎝⎜⎞

⎠⎟σσ

ττ

32

≤≤ 1

σ ρeq Ed ydf, min ·≤

λσ

λσ

λττx p

y

x crz p

y

z crp

y

cr

f f f,

,,

,,; ;

·= = =

3

B. Braun/U. Kuhlmann · Reduced stress design of plates under biaxial compression



In the next step, separate reduction factors are determinedfor each slenderness:

(7)

where:κx reduction factor for longitudinal stresses, interpolated

value κx,PK where column-like buckling is relevantκz reduction factor for transverse stresses, interpolated

value κz,PK where column-like buckling is relevantκτ reduction factor for shear stresses

Finally, the general verification format to account for in-teraction between stress components according to Eq. (8)is based on an interaction equation with the parametersei =1, 2, 3 and V:

(8)

where:σEd; τEd design load valuesσx,Rd = κx · fyd; design resistance valueσz,Rd = κz · fyd; design resistance valueτRd = κτ · fyd; design shear resistance valuefyd design yield strengthe1 = 1 + κ4

xe2 = 1 + κ4

ze3 = 1 + κx · κz · κ2

τV = (κx · κz)6 when σx,Ed and σz,Ed are both compres-

sion, otherwise

As the global slenderness was introduced in EN 1993-1-5,so the concept of the reduced stress method in DIN 18800-3was varied. In DIN 18800-3 the interaction verification isonly allocated to Eq. (8), for which parameters ei = 1, 2, 3and V have been calibrated to fit the database in [8]. Incontrast, the interaction in EN 1993-1-5 is taken into ac-count both at elastic critical load level and finally by theinteraction equation according to Eq. (4). Thus, it is notpossible to transfer simply the values of the parameters de-termined for the concept in DIN 18800-3 to the conceptin EN 1993-1-5.

3 Plates under biaxial compression3.1 General

Figs. 1 and 2 illustrate the procedures described in section2.2 for square plates under uniform biaxial compression.Explanatory notation is given in Fig. 3. It can be seen thatthe global slenderness concept in EN 1993-1-5 leads to avarying slenderness for a given b/t ratio, thus taking inter-action into account in a very elaborate way. In contrast, theslenderness according to DIN 18800-3 is based on separatestress components and thus is constant for a given b/t ra-tio. The projection of the interaction curves onto the ρx–ρzplane is compared in Figs. 4 and 5 for representative b/tratios. It can be shown that the curves are significantly dif-

κ λ κ λ κ λτ τx x p z z p p, , ,; ;( ) ( ) ( )

σ

σ

σ

σ

σx Ed

x Rd

e

z Ed

z Rd

e

xV

,

,

,

,

,·

⎛

⎝⎜⎜

⎞

⎠⎟⎟

+⎛

⎝⎜⎜

⎞

⎠⎟⎟

−1 2

EEd z Ed

x Rd z Rd

ed

Rd

e·

··

,

, ,

σ

σ σττ

⎛

⎝⎜⎜

⎞

⎠⎟⎟

+⎛

⎝⎜⎞

⎠⎟≤3 1

3

V x Ed z Ed

x Ed z Ed

=σ σ

σ σ, ,

, ,

·

·

ferent. But this does not depend on the reference strengthsat the axes; rather, it depends on the capability of the inter-action surface, which is defined by the format of the un-derlying interaction equation, see also [7, 13]. In fact, whenthe slenderness approaches the border to class 3 cross-sec-tions, i.e. b/t ≈ 30 here, a stability-induced reduction is ex-pected under biaxial compression, which is covered byEN 1993-1-5 but not by DIN 18800-3. On the other hand,a linear interaction should be approached with increasing


Fig. 1. Interaction scheme according to EN 1993-1-5, chapter10 (α = 1, fy = 355 N/mm2)

Fig. 2. Interaction scheme according to DIN 18800-3 (α = 1,fy = 355 N/mm2)

Fig. 3. Notation for a plate subjected to biaxial compression


slenderness – covered by DIN 18800-3 but not by EN1993-1-5. In principle, the interaction equation should ba-sically reflect the mechanical behaviour irrespective of theverification format. Since the EN 1993-1-5 concept consid-ers the full stress field, it is better suited to describing themutual influence of stresses. However, the non-negligiblediscrepancies for high slenderness require clarification.

3.2 Consequences of earlier work

Earlier research work comprises a large number of studies,though few experimental ones, which date back to the 1970sand 1980s. Later work mainly focused on summarizing andre-evaluating these studies. The evaluation revealed a largeknowledge base where, however, the large number of in-fluential parameters always led to a discussion about thequality and usability of the results. For this reason, addi-tional simulations were carried out in [4] which have beenused for reassessing earlier work and which form the prin-cipal basis for the proposal in section 4. Some aspects re-lated to the parameter study of biaxially compressed platesare presented in the following.

3.3 Effect of imperfections and panel aspect ratio

For a rectangular plate, the imperfection shape that leadsto the smallest resistance depends on the number of half-


waves and the corresponding loading direction. Figs. 6aand 6b show two modal imperfection shapes of a biaxiallyloaded plate with a panel aspect ratio of α = 3. For each ofthe uniaxial loadings, the corresponding mode leads tothe smallest resistance. However, the one-half-wave modeincreases the resistance when the short edges are mainlyloaded because the out-of-plane deformation usually pre-vents the plate from switching to its natural three-half-wave mode. On the other hand, in the case of biaxial com-pression, the inflexion points of the three-half-wave modeact like a stiffening of the plate (thick lines in Fig. 6b).

In the parameter study, plates with a panel aspect ra-tio of α = 1 and 3 were considered. A large number of ini-tial mode shape combinations was studied in [5] for panelaspect ratios of α = 3 but a smaller failure resistance thanthat achieved with the three- and one-half-wave mode shapescould not be produced. Measurements of the initial imper-fection shapes occurring in normally fabricated plates [6]support that, in principle, the one-half-wave mode shapeis the one most likely to occur in practice. Thus, the three-half-wave mode shape is usually on the safe side when theloading of the short edges begins to dominate. The inter-action curve shape in Fig. 6c can therefore be seen as anexample of a lower bound. In terms of imperfections, thesquare plate is the simplest one, albeit the worst case, be-cause the imperfection shapes for each stress componentusually coincide.


a) b/t = 30 b) b/t = 65

Fig. 4. Comparison of interaction curves for biaxial loading (α = 1, fy = 355 N/mm2)

a) b/t = 30 b) b/t = 45

Fig. 5. Comparison of interaction curves for biaxial loading (α = 3, fy = 355 N/mm2)



The imperfection amplitude has a pronounced influ-ence in the small to medium slenderness range. In the para-meter study, equivalent geometric imperfections were usedwith amplitude values of min(a/200; b/200) according toEN 1993-1-5, Annex C, and min(a/420; b/420) based on [12].It was shown that recalculating buckling curves succeedsonly with the reduced amplitude, see also [9]. Residualstresses were studied but it was shown that they do not in-crease the quality of the parameter study significantly.

3.4 Effect of edge boundary conditions

The edge boundary condition can be subdivided into in-plane and rotational restraints. Extreme in-plane edge con-figurations are the “unconstrained” case (free to move in-plane, stress-controlled) and the “constrained” case (forcedto remain straight, strain-controlled). Examples of the con-figurations and their effect on stress distributions are shownin Figs. 7a and 7b for uniaxially loaded plates. Fig. 7c illus-trates the same situation for all edges constrained and bi-axial compression. The case with unconstrained edges andbiaxial compression was also addressed in the parameterstudy. It is concluded from the comprehensive recalcula-

tions of buckling curves in [4] that an appropriate assess-ment of the reference strength is required in order to achievemeaningful results. Thus, in the parameter study, the buck-ling curves chosen represent the corresponding edge bound-ary condition as precisely as possible. EN 1993-1-5, Annex B,with αp = 0.34 and λ–p0 = 0.70, is used for the case with alledges unconstrained. In the case of all edges constrained,EN 1993-1-5 does not provide an adequate reduction func-tion, so for this case the authors’ own databased bucklingcurve was used.

Rotational edge boundary conditions are usually cov-ered at elastic critical load level since their effectivenessdoes not depend on large deflections.

It was found that an appropriate choice of bucklingcurve facilitates the use of few interaction curves whichare then able to cover different edge boundary conditions.

4 Proposal for an improved design rule

In order to achieve an adequate description of the stabilitybehaviour, a modification of the interaction equation inEN 1993-1-5, chapter 10, is required. The proposal focuseson the high slenderness range where, for square plates, a


a) one-half-wave mode b) three-half-wave mode c) effect on interaction curve

Fig. 6. Comparison of imperfection shapes (α = 3, fy = 355 N/mm2)

a) at ultimate load, unloaded edges b) at ultimate load, unloaded edges c) at ultimate load, all edges con-unconstrained constrained strained and biaxially loaded

Fig. 7. Membrane stress distributions, loaded edges constrained (b/t = 100, fy = 355 N/mm2)


linear interaction relationship should be approached. Forrectangular plates, the favourable imperfection mode shapesshould be accounted for. In all cases, a tunable interactionequation has to be formulated. A generalized form of suchan equation based on the von Mises yield criterion has al-ready been given in Eq. (8), taken from DIN 18800-3. Forbiaxially compressed plates, the most relevant parameterwith an effect on the interaction curve is the factor V.Thus, a tunable interaction equation can be written in sim-plified form:

(9)

where:σEd; τEd design load valuesσRd; τRd design resistance valueV factor

The approach for the factor V is made according to Eq. (10):

(10)

where:ρx; ρz reduction factorse tunable exponent

In the first step, the results of the parameter study on squareplates are used to study the effect of the factor V. Due tothe symmetry of the square plate, the curvature can be de-fined by points located on the bisecting plane at an angleof 45°, see Fig. 8. The comparison of simulations, designrules and proposals in this plane is given in Fig. 9. A dis-tinction is made between unconstrained and constrained in-plane edge restraints. The Winter curve from EN 1993-1-5,section 4.4(2), is provided as a reference, but it should not

σσ

σσ

σx Ed

x Rd

z Ed

z Rd

x EdV,

,

,

,

,·⎛

⎝⎜

⎞

⎠⎟ +

⎛

⎝⎜

⎞

⎠⎟ −

2 2··

·,

, ,

σσ σ

z Ed

x Rd z Rd

⎛

⎝⎜

⎞

⎠⎟ ≤ 1

V x ze

= ( )ρ ρ·


be considered further because it represents loaded edgesconstrained and unloaded edges unconstrained, which isnot applicable to the biaxial load scenarios here.

Summing up, the tunable parameter e becomes 1.7when fitted to the data as a lower bound. It is governed bythe medium slenderness range and becomes slightly con-servative for high slenderness values. Bearing in mind thatthe recalculation of the buckling curves in the medium


Fig. 8. Plane utilized for calibrating the exponent e in Eq. (10)

Fig. 9. Calibration of V and effect of in-plane edge restraints(α = 1, all edges hinged)

a) all edges unconstrained

b) all edges constrained



slenderness range is only successful with a reduced imper-fection amplitude of b/420, see section 3.3, a similar rea-soning applies to the calibration of the interaction equa-tion, and it is shown that e can be simplified to 1.0. Basedon this, it is proposed to modify Eq. (10.5) in section 10(5b), EN 1993-1-5, according to Eq. (11). In order to im-prove the layout of the verification procedure it is pro-posed to delete sections (2) and (5a):

(11)

where:σEd; τEd design load valuesσx,Rd = ρx · fyd; design resistance valueσz,Rd = ρz · fyd; design resistance valueτRd = χw · fyd; design shear resistance valuefyd design yield strengthV = ρx · ρz when σx,Ed and σz,Ed are both compres-

sion, otherwise V = 1

Comparing the simulations and the experimental resultswith the proposal reveals sufficiently good agreement. Eq.

σσ

σσ

σx Ed

x Rd

z Ed

z Rd

x EdV,

,

,

,

,·⎛

⎝⎜

⎞

⎠⎟ +

⎛

⎝⎜

⎞

⎠⎟ −

2 2

σσσσ

ττx Rd

z Ed

z Rd

Ed

Rd,

,

,

· ·⎛

⎝⎜

⎞

⎠⎟

⎛

⎝⎜

⎞

⎠⎟ +

⎛

⎝⎜⎞

⎠3 ⎟⎟ ≤

2

1

(11) is compared with EN 1993-1-5 and DIN 18800-3 inFigs. 10 and 11. For square plates, it can be shown that a lin-ear interaction is approached as the plate becomes moreslender. As a result, the proposal more or less coincides withDIN 18800-3. However, for a small slenderness, a strongerinteraction has to be taken into account so that the pro-posed interaction equation becomes more conservative,which is supported by the simulations. For panel aspect ra-tios other than square, the imperfection shape was proved tohave a significant influence on the curvature of the interac-tion curves. DIN 18800-3 is too conservative, whereas EN1993-1-5 is the opposite, i.e. too favourable. Here, the pro-posed interaction equation is able to take the positive influ-ence of the imperfection shape successfully into account.

5 Conclusions and implications for further work

As the global slenderness was introduced in EN 1993-1-5, sothe concept of the reduced stress method in DIN 18800-3was varied. The basic idea behind EN 1993-1-5 follows for-mer German guideline DASt-Ri 012. The new concept offersadvantages over the today’s numerical procedures whichallow the elastic critical load factor of the full stress field tobe determined in a single step, thus taking interaction intoaccount in a very elaborate way. However, it is shown in this


a) b/t = 30 b) b/t = 65

Fig. 10. Comparison of interaction rules and proposal (α = 1, fy = 355 N/mm2)

a) b/t = 30 b) b/t = 45

Fig. 11. Comparison of interaction rules and proposal (α = 3, fy = 355 N/mm2)


article that the interaction verification in its pure formatbased on the von Mises yield criterion is not able to repre-sent the actual mechanical behaviour of biaxially compressedplates. The proposed modification leads to appropriate andplausible results without over-complicating the interactionequation.

A recent publication [10] shows that a further improve-ment of EN 1993-1-5, chapter 10, has been facilitated basedon [4].

Besider that, when the loading changes from com-pression–compression to compression–tension, the globalslenderness concept is able to consider the effect of tensilestresses, see Figs. 4 and 5. However, only few pilot studieshave looked at this effect. Since tensile stresses can havenot only a stabilizing effect on buckling but also a plasticdestabilizing influence, further work in this field will benecessary in the future.

References

[1] DASt-Ri 012. Beulsicherheitsnachweise für Platten, 1980.[2] DIN 18800. Steel Structures – Part 3: Stability, Buckling of

plates, 1990 and 2008 (in German).[3] EN 1993-1-5. Eurocode 3: Design of steel structures – Part

1–5: Plated Structural Elements, 2007.[4] Braun, B.: Stability of steel plates under combined loading.

Dissertation, No. 2010-3, Institute of Structural Design, Uni-versity of Stuttgart, 2010 (http://elib.uni-stuttgart.de/opus/volltexte/2010/5887).

[5] Dowling, P. S., Harding, J. E., Slatford, J. E.: Strength ofShips Plating – Plates in biaxial compression. CESLIC Re-port SP4, Department of Civil Engineering, Imperial College,London, 1979.

[6] Fischer, M., Grube, R., Rieger, H., Wenk, P.: Messungen derVorverformungen von beulgefährdeten Stahlblechen mit undohne Steifen. University of Dortmund, 1989.


[7] Johansson, B., Veljkovic, M.: Review of plate buckling rulesin EN 1993-1-5. Steel Construction – Design and Research 2(2009), No. 4, pp. 228–234.

[8] Lindner, J., Habermann, W.: Zur mehrachsigen Bean spru -chung beim Plattenbeulen. Festschrift Prof. Joachim Scheer,1987, pp. 159–183.

[9] Rusch, A., Lindner, J.: Load-carrying capacity of thin-walledcross-section elements in view of imperfections. Stahlbau 70(2001), No. 10, pp. 765–774 (in German).

[10] Maur, J., Schmidt, H., Verwiebe, C.: Buckling stress designcheck for plane and curved steel surface structures undercombined membrane stressing – a comparative analysis of theEurocode formats. Stahlbau 80 (2011), No. 11, pp. 804–813(in German).

[11] Naumes, J., Feldmann, M., Sedlacek, G.: Demonstrationof the common basis of method 1 (effective width approach)and method 2 (stress limit approach) for the plate-bucklingassessment of built-up steel components according to Euro -code 3 – Part 1–5. Stahlbau 78 (2009), No. 3, pp. 139–147 (inGerman).

[12] Usami, T.: Effective width of locally buckled plates in com-pression and bending. Journal of Structural Engineering 119(1993), No. 5, pp. 1358–1373.

[13] Winterstetter, T.: Buckling design of steel plate and shellstructures under combined loading. Stahlbau 71 (2002), No. 11,pp. 816–822 (in German).

Keywords: stability; plate buckling; interaction; biaxial com-pression

Authors:Dr. Benjamin Braun: Federal Waterways Engineering & Research Institute, Kussmaulstrasse 17,76187 Karlsruhe, GermanyProf. Dr. Ulrike Kuhlmann: Institute of Structural Design, University of Stuttgart, Pfaffenwaldring 7, 70569 Stuttgart, Germany


chain caused steel demand to weaken inthe 2nd half of 2011.

Early 2012, reduced domestic andimport supply appears to be balancedwith softer demand. Rather low stock levels – particularly at distributors – willrequire some replenishment in H1-2012,which will ease later in the year. Buyersanticipating changes in prices and supplycould exacerbate this seasonal pattern,despite the fact that inventories will con-tinue to be managed cautiously.

Eurofer director-general GordonMoffat sees the market strengtheningagain in 2013: “The market will take apause in 2012, but we foresee only avery mild reduction in demand. Realand apparent steel consumption willstrengthen again in 2013”.

Please visit www.eurofer.org for moreinformation.

Economic and Steel Market Outlook2012–2013

Eurofer’s Q1-2012 economic and steelmarket outlook shows that while the EUeconomy probably slipped into a shallowrecession in the final quarter of last year,activity of downstream steel-users hasremained rather firm.

The latest economic data confirm thatthe Eurozone’s debt crisis is eating its wayinto the real economy. Since mid-2011,economic sentiment has come down shar-ply on the back of the further intensifica-tion of the sovereign debt crisis. Risk aver-sion in the financial sector has led to acontraction in credit supply. While it iswidely recognised that austerity is requi-red to avoid the sovereign debt crisisfrom getting worse, it will also stifle eco-nomic growth and affect employment.Investment and private consumption

News

will be subdued until confidence returnsto the markets.

Eurofer director-general GordonMoffat: “When confidence will return to the markets is still difficult to say, butit is encouraging that forward lookingindicators such as industrial confidenceand manufacturing PMIs appear to havetaken a turn for the better in recentweeks. Companies will remain cautiousfor the time being, but there is no indi-cation that industrial orders will fall offa cliff. Order backlogs are quite strong.The outlook for most steel-users is rela -tively benign”.

EU apparent steel consumption grew7.2 % in 2011. Steel purchasers waitingat the side-lines after the summer due torising concerns about the economic situa-tion, downstream business conditionsand access to financing and credit to-gether with destocking in the distribution



1 Introduction

For the sciences, application-orientedcooperative research (CR) is an im-portant element of their interactionwith the economy. It also has a posi-tive effect on the innovation behav-iour of the companies participating [1].The benefits of pre-competitive formsof CR for small and medium-sized enterprises (SMEs) in particular havebeen clearly demonstrated [2]. Thesteel industry in Germany supportsthis kind of application-oriented re-search by engaging in joint pre-com-petitive research projects in coopera-tion with scientific institutions andcompanies from various steel applica-tion sectors. The Research Associa-tion for Steel Application (FOSTA) isresponsible for managing these pro-jects on behalf of the steel industry inGermany. The projects pursued by thisbody produce application-oriented re-sults in the interests of all participants.Projects in cooperation with the con-struction sector account for a signifi-cant proportion of this type of steelapplication research. The startingpoint in the research management ofthese projects is described in [3] on

the basis of typical innovation attrib-utes in the construction sector andcharacteristic deficiencies in the ex-isting CR system. These aspects arereproduced in summarized form be-low as the basis for further examina-tion, and are partly expanded upon.The subsequent discussion on the sta-tistical results of a survey devoted tothe project expectations of scienceand industry in a FOSTA researchcluster entitled “Sustainability of steelin the construction sector (NASTA)”enables the development of the taskprofile for an innovation broker (IB)operating within the framework of in-tegrated management of CR betweenthe steel industry, the constructionsector and research institutes [4].

2 Summary of the innovation characte-ristics of the construction sector andthe existing system deficiencies incooperative research

The construction sector is often char-acterized as being not very innovativeand having a traditional structure [5].R&D investment and the number ofregistered patents are both extremelylow in relation to other sectors [6, 7].Innovations occur less within theframework of the sector’s own inno-vation projects and more in response

to the special demands of specific con-struction projects [8, 9]. Innovationopportunities therefore also dependon the many regulations that definespecifications for materials, technolo-gies and processes [6, 9]. Consequently,the industry itself often finds itself ina “locked system”, from which it isdifficult to break away [5]. In the lightof this, the construction sector is of-ten seen not primarily as a developerof technologies, but rather in the roleof a user of technologies that haveoriginated in innovative companies inthe supply chain. According to [10],the industry structure represents anadditional challenge to achieving in-novation success in the constructionsector. This structure is characterizedby SMEs that do “not (often) havethe resources and the knowledge todevelop innovations on their own”[11], that “(often) have an organiza-tional slack in comparison with largeconstruction firms that invest in strate-gies with long-term return” [12], andthat “usually invest in their own R&Donly to a small degree” [13]. The steelindustry in Germany is a key supplierof materials and products for the con-struction sector [14] and therefore ac-tively tries to generate new innova-tions for steel applications in the con-struction industry within the scope ofapplication-oriented pre-competitiveCR with the construction sector. Theresulting innovation process can bedivided into four phases – early, ad-ministrative, project and late – asshown in Fig. 1.

These are in turn subdivided intofurther phases that essentially depictthe individual steps of the innovationprocess in CR. This involves, in par-ticular, the search for suitable re-search ideas of interest to the steel

Gregor Nuesse*Mukesh LimbachiyaRoland HerrAlan Ellis

Managing interdisciplinary applied researchon sustainability in construction with the helpof an innovation broker

The article reports on how the steel industry in Germany, as a key supplier of materialsand innovations to the construction sector, is carrying out a comprehensive optimizationof the management of pre-competitive cooperative research (CR) on steel applications inthe construction industry. A significant aspect of this is the central use of an innovationbroker (IB) role throughout the innovation process. The varying project expectations ofscience and industry and the related IB tasks are identified as part of an interdisciplinaryresearch cluster on the sustainability of steel in the construction industry. A task profileis drawn up for the IB in cooperative research on the basis of various key person con-cepts from industrial innovation research. It is expected that the consistent implementa-tion of this profile will increase the innovation successes from CR between the steel in-dustry and the construction sector.

Articles

DOI: 10.1002/stco.201200006


11_041-052_Nuesse (SC8)_000-000_Steel Constr. (3sp).qxd 09.02.12 12:33 Seite 41

42

G. Nuesse/M. Limbachiya/R. Herr/A. Ellis · Managing interdisciplinary applied research on sustainability in construction with the help of an innovation broker

Steel Construction 5 (2012), No. 1

and construction industries (Initia-tive) and the subsequent further de-velopment of these ideas until a re-search proposal is ready for submis-sion, including the finding of suitableindustrial and academic project part-ners (Develop). This is followed bythe organizing of funding possibilitiesfrom industry and the public sector(Finance), early agreement on the useof the potential results (Own) and theactual research process undertakenby research institutes (Research). Thesubsequent initial application of theresearch results achieved in coopera-tion with industrial application part-ners (Use) and the possible transfer toand use of the research results ob-tained in other possible industrial steelapplication sectors (Transfer) concludethe process. Currently, CR in practicesometimes means passing through theindividual stages completely, but oftenonly partially. In addition, in order tosurmount possible gates between thephases, there are various players, suchas universities, research funders orpartners from industry, who often op-erate independently of each other.This makes the whole CR innovationprocess a complex and sometimes ex-tremely fragmented system. As a logi-cal consequence, “deficiencies” oftenarise in the course of the process. Inthe light of this, Sexton et al. [11] note,for instance, on relevant CR projectsthat “unsuccessful innovation oftenhappens because of a lack of marketfocus (in the project)”. Barnes et al.[15] are critical of the fact that “theobjectives (of a project) were often notspecific enough with respect to theexpected outcomes and were largelyopen to interpretation in terms”; theyalso stress that “there have been in-evitable differences in the require-ments and expectations of the (pro-ject) partners …” For the projectphase, Abbott et al. [16] comment that“it was felt that at times universitiesand industry talked different lan-guages”. Couchman et al. [17] addthat “universities are (often) not very

effective in commercializing newideas” and that “the knowledge uni-versities produce is not (often) in aform that can be directly applied orcommercially exploited.” Sexton et al.[11] argue that “current technologytransfer mechanisms are not suffi-ciently informed by, or engaged with,company strategic direction and orga-nizational capabilities”. FOSTA wouldlike to counteract these possible defi-ciencies through the central use of anIB role while seeking to improve theinnovation process. This approach isdescribed in greater detail below.

3 Approach to optimizing the manage-ment of CR between the steel andconstruction industries

The steel industry in Germany, as al-ready mentioned at the start, launchedan interdisciplinary FOSTA researchcluster entitled “Sustainability of steelin the construction sector (NASTA)”in early 2011 as part of its CR withthe construction industry. This re-search cluster, comprising six subpro-jects in total, is engaged in develop-ing indicators for the quantitative assessment of sustainability in steelconstruction. There are, in total, 30research institutes from 11 differentdisciplines and 200 industry expertsinvolved in the NASTA project [18].The central role of an innovation bro-ker (IB) was established within theproject as part of an integrated ap-proach to innovation management inCR. The aim of the IB is to take ac-count of the overriding requirementsin the management of interdiscipli-nary tasks set but also, in particular,to ensure that the special features ofthe construction sector in the innova-tion process are considered whilehelping to prevent the deficiencies inCR outlined earlier. Publications oninnovation management contain vari-ous, mostly company-related, key per-son concepts that can be used to cre-ate a possible task profile for the IB.The champion, the promoter and the

gatekeeper models can be called thechief markers pointing the way. Therespective characteristics of these mod-els are summarized in Table 1.

So far, the concepts cited inTable 1 have been given only scantconsideration in the majority of theproposals regarding the establishmentof an IB role in construction industryCR. The identification of an IB is of-ten proposed only in the project team,i.e. at the research stage, or, for exam-ple, as a key administrative positionin the management of a research in-stitution, without any bearing on aspecific project or subject. The tasksare then described only in generalform without specific reference to thewhole innovation process (see, for ex-ample, [16, 33]). The question there-fore arises as to how the IB role canbest be integrated for CR in the in-novation process and how an IB’stask profile can be purposefullystructured to meet CR requirements.In this context, the questions for theCR under examination here also re-main: What interests are the partic-ipants from science and industrypursuing through their participa-tion in the project? And: How canthese interests be considered appro-priately in the project with the IB’shelp? With this in mind, so-called in-dustry working groups (IWGs) at-tached to the respective subprojectshave been established in the NASTAproject in addition to the IB role,their aim being to provide the innova-tion project with specialist support.These IWGs do not consist of a lim-ited number of selected companies,as is often the case, but rather areopen to innovation and are accessibleto all interested industry representa-tives without restrictions.

4 Survey of project participants: formu-lation of questions, structure of eva-luation, approach to interpretation

In order to answer the questionsposed earlier regarding the structur-

Fig. 1. Phase model in cooperative research (CR) on the basis of [4]



ing and integration of the IB role inCR and regarding the management ofthe varying interests in the innova-tion process within IWGs, a survey ofthe heads of research institutes (n =21) and various industry representa-tives (n = 50) from the IWGs was con-ducted as part of the NASTA project.This was carried out within the firstIWG meetings held in late 2010 andearly 2011 once the early and admin-istrative innovation phases of CR hadbeen completed and the project vol-ume amounting to € 6 million hadbeen approved. The survey was con-ducted with standardized question-

naires based on the following ap-proaches:

1. The research institutes were ques-tioned about their expectations ofthe forthcoming IWGs. They werealso asked for a retrospective as-sessment of the possible IB tasksfor the early and administrative innovation phases of CR alreadycompleted.

2. The industry representatives wereasked to give the reasons that re-sulted in their support for the re-spective NASTA subproject throughtheir participation in the IWG.

3. Both groups were asked abouttheir expectations of the IB’s rolein the forthcoming project phaseand in the late innovation phasesof CR between science and indus-try.

For the replies to approaches 1 and 2,a four-point Likert scale was selectedwith which the possible expectationsof the research institutes in relationto the IWGs as listed in the question-naire, the assessment of the possibleIB tasks in the early and administra-tive innovation phases and the possi-ble reasons for industry’s participa-tion in the IWGs could be assessed.The scale provides for an assessmentof (1) for “unimportant” to (4) for“very important”. For the third ap-proach, which was directed at bothgroups in the same form, a four-pointLikert scale was again used, enablingthe science and industry representa-tives polled to assess the possibletasks of the IB role in the projectphase and the late innovation phases.The scale provides for a range of re-sponses from (1) for “undesirable” to(4) for “very desirable”. If the respon-dents did not wish to reply to certainquestions, they were asked to leavethe scale empty. In selecting the re-spondents, care was taken to ensurethat the disciplines included in theNASTA project should also be repre-sented in the survey in accordancewith their project participation. Amongthe research institutes, this includedthe fields of structural steel engineer-ing/lightweight steel construction/composite construction (43 %), archi-tecture/building services/structural de-sign (24 %), life cycle management(19 %) and construction management/business administration/work organi-zation (14 %). Among the industry rep-resentatives, experts were questionedfrom engineering firms for construc-tion planning and construction pro-ject management (30 %), from steelproduction (20 %), from steel con-struction (20 %) and from comple-mentary fields (30 %), e.g. manufac-turers of building products and con-struction software, industrial associa-tions. As part of the statistical analy-sis, the following evaluation steps wereperformed using the PASW Statistics18 program for the replies from theresearch institutes and those from in-dustry separately.


The gatekeeper model [28], [29], [30]

– Project-independent person model whose main task is to pass on and obtain spe-cialist knowledge that transcends organizational boundaries.

– The gatekeeper possesses extensive and up-to-date expertise and is also a networkerand communicator.

– This person performs a prominent role in interface management and is the personwhose specialist advice is in demand.

– Whereas the relationship promoter tends to be more active in the exploitation ofinnovations, the gatekeeper imports knowledge for an innovation process.

The (product) champion model [19], [20], [21], [22]

– One-person concept: an especially dedicated person is driven by their own traitsand convictions to push through an innovation in a company against all odds.

– The champion possesses technical expertise, detailed company and market know-ledge, drive and political skill.

– That person takes charge of the innovation at the conception stage, frees the inno-vators from the rules of the organization, convinces others to support the innova-tion, provides the necessary resources, protects the innovator against disruptionand pushes the innovation forward.

– There is no more precise description of the champion’s role.

– Its impact on success has been repeatedly demonstrated but the focus on one person can also have negative effects very quickly.

The promoter model [23], [24], [25], [26], [27]

– Departure from the one-person focus.

– Theory-based and empirically tested concept.

– For the bridging of barriers to participation within an organization, it includes a“power promoter” for those not wanting to take part and a “know-how promoter”for those lacking the necessary specialist knowledge.

– The know-how promoter possesses expert knowledge and is often the nucleus ofthe innovation.

– The power promoter has high hierarchical potential and ensures the resources required and the necessary decision-making for the project’s preservation.

– The process promoter, usually from middle management, interlinks both roles, accompanies the innovation process, motivates and coordinates within the process.

– A relationship promoter is added in order to bridge the inter-company barrier of alack of knowledge among external partners.

– This promoter has a network of personal contacts, both inside and outside thecompany, and brings together people from various institutions.

– As yet, there has been little research into the motives governing the actions of promoters.

Table 1. Summary of varying key person concepts in the innovation process inline with [31] and [32]


Step 1:For the purpose of creating a rankingwithin the individual reply optionsfor a question, the mean of the re-spective assessments was determinedat the start. In order to improve themeaningfulness of the mean values,the related 95 % confidence inter-vals (CI) were also established. More-over, a confidence belt with corre-sponding upper and lower limits wasdetermined for the mean value with adefined probability depending on sam-ple size and respective standard devi-ations.

Step 2:As a supplementary aspect for illus-trating priorities within the reply op-tions, the percentages of the personsquestioned who favoured a specificaspect of the IWG and at the sametime a specific IB task were also as-certained. To this end, positions (1)and (2) of the respective four-pointLikert scales were summarized withindata transformation and interpretedwith a somewhat negative slant. Inthe same way, positions (3) and (4) ofthe respective Likert scales weregrouped together and interpreted witha somewhat positive slant. Generat-ing cross-tables from these data be-tween the reply options for IWG as-sessment and IB tasks means it ispossible to identify aspects that havebeen simultaneously assessed as rela-tively favourable. A summarized viewof the respondents to both questions(IWG and IB) is thus produced at thesame time.

Step 3:In order to identify possible trend-forming connections between indi-vidual expectations of the IWG andindividual IB tasks, Kendall’s Tau-bwas then determined as the rank cor-relation coefficient for all possiblecombinations of the aspects priori-tized previously. In the process, theranking of the existing replies is com-pared in each case with an expecta-tion of the IWG and an IB task anddisplayed in the form of a correlationcoefficient. However, the chief infor-mative value of the correlation deter-mined is only expressed through thesignificance level ascertained at thesame time. Therefore, for the purposeof this analysis, only those correla-tions with a trend-forming connec-


tion defined as significant (95 %) orhighly significant (99 %) were inter-preted.

5 Statistical analysis5.1 Expectations of heads of research

institutes regarding the IWG ac-companying the project and assess -ment of the corresponding IB tasks

Within the survey of the heads of re-search institute, the four-point Likertscale already described was used atthe start to enquire about the variousexpectations of an IWG. The left-handcolumn of Table 2 shows results indescending order on the basis of themean values determined empirically.The upper and lower limits of the re-lated 95 % confidence intervals (CI)are also shown. Based on the samesystem of classification, the right-handcolumn of Table 2 shows the resultsof the assessment of possible IB tasksby the heads of research institutes,again with a four-point Likert scale.

The mean values for what theheads of research institutes expectfrom IWGs range from 2.7 to 3.2,with the related 95 % CIs varyingfrom 2.2 to 3.6. The various aspectsare assessed favourably in general, al-though it is noticeable that from the8th aspect onwards, the mean forIWG expectations falls short of 3.0and the lower limit of the 95 % CIdrops to 2.4 between the 7th and 8thaspects. The individual assessmentsof the subsequent aspects 9 and 10decline correspondingly further andan increase in the corresponding CIscan be seen.

In the assessment of possible IBtasks by the research institute headsquestioned, as shown in the right-hand column of Table 2, the meanvalues range from 2.8 to 3.8, with thecorresponding 95 % CIs varying from2.3 to 4.0. Here, too, there is a gener-ally favourable assessment of possibleIB tasks, although an increase in the95 % CI owing to the drop in thelower limit from 2.9 to 2.7 from IBtask 10 onwards is striking. For thefirst time, the mean value of 3.0 is notreached from IB task 13 onwards.

Since the aim is to develop an IBtask profile from the survey on thebasis of the expectations of the vari-ous participants in the IWG accom-panying the project, those IWG as-pects and IB tasks that were given a

simultaneous assessment tending to-wards the favourable by the respon-dents were identified in the secondevaluation stage. For the research in-stitute heads, the necessary data trans-formation and cross-table calculationproduced a simultaneously favourableassessment of one of the first six IWGaspects and of the whole of the first10 IB tasks by 70 to 90 % of the re-spondents. This means that well overtwo-thirds of those questioned ap-proved of the first six IWG aspectsand the first 10 IB tasks at the sametime. For the 7th IWG expectation, aremaining simultaneously favourableassessment could be established atthis level (for exactly 76 % of respon-dents) with the first four IB tasks. Asimultaneously favourable assessmentin the range of 23 to 66 % of respon-dents was demonstrated for all otherIWG aspects and IB tasks. For fur-ther analysis, in the light of the broadsimultaneous approval by the researchinstitute heads, the first seven IWGexpectations are interpreted as im-portant and the first 10 IB tasks asdesirable. This interpretation is alsoconfirmed by the mean values and95 % CIs given above.

The aim of the third stage of theevaluation is to determine the rankcorrelations verifiable by calculationbetween the aspects interpreted asimportant and desirable and to usethis to generate trend-forming asser-tions regarding a possible connectionbetween key IWG expectations anddesirable IB tasks. Table 3, which in-dicates the respective rank correla-tion coefficient and the related signif-icance level, summarizes the aspectsidentified. With the coefficients ascer-tained, at between 0.30 and 0.49, amoderately strong positive connec-tion can be assumed between therankings from the Likert scales usedto assess IWG expectations and IBtasks. The coefficients between 0.50and 0.69 indicate a strong positiveconnection between the aspects citedin each case.

In order to complete the wholepicture in a further analysis, a briefoverview of the IB tasks for the earlyand administrative innovation phases,as already described in [3], i.e. for theperiod before the project starts, willnow be given: the retrospective as-sessment of possible IB tasks in theearly and administrative innovation




phases by the heads of research insti-tutes produces mean values of 3.2 to3.9 from the survey. The 95 % CIrange of mean values varies from 2.9to 4.0. Overall, the research instituteheads therefore assess the use of anIB as desirable, even for the early andadministrative innovation phases of

CR. In particular, the early brokeringof industry contacts (3.0 | 3.3 | 3.6),the central acceptance and collectionof outlines for project ideas (2.9 | 3.2 |3.5) and the organization of pre-eval-uation discussions with industry (3.2 |3.4 | 3.6) are especially highlighted asIB tasks. In the administrative phases,

the IB should, in the view of the re-search institute heads, in particularinitiate and, if possible, also conductcommunication with prospective fund-ing providers (3.7 | 3.9 | 4.0). Shouldconflicts of interest arise during theearly and administrative innovationphases, the IB should, in the view of


Research institutes heads’ expectations of the industry working groups (IWG)

(Lower limit of the 95 % CI | Mean | Upper limit of the 95 % CI)

1.)IWGs offer the opportunity to give the research project a

practical orientation (3.0 | 3.2 | 3.6)

2.)IWGs offer us and the representatives from industry the opportunity for joint know-how development on the

research topic (2.9 | 3.2 | 3.5)

3.)IWGs represent milestones for the scientific staff over the project period at which new interim results have to

be presented (2.9 | 3.2 | 3.5)

4.)IWGs offer the opportunity to gain access to

topic-related information which has not yet been considered (2.9 | 3.2 | 3.5)

5.)IWGs are important to us in initiating new projects with

industry involvement (2.9 | 3.1 | 3.5)

6.)IWGs enable a practice-oriented evaluation of the

relevant research progress (2.9 | 3.1 | 3.4)

7.)IWGs offer the opportunity to present the activities of

our research institute to industry (2.7 | 3.0 | 3.4)

8.)IWGs are important to scientific staff for collecting

contacts in industry (2.4 | 2.9 | 3.3)

9.)IWGs lead during the project period to the creation of new subtasks which help to prepare for industrial

application (2.5 | 2.8 | 3.1)

10.)IWGs are important in initiating consulting services

through us for companies (2.2 | 2.7 | 3.0)

Research institutes heads’ views of the tasks of the innovation broker (IB) in the project

(Lower limit of 95 % CI | Mean | Upper limit of 95 % CI)

1.) Publishing the final report (3.6 | 3.8 | 4.0)

2.)Organizing events with the relevant target audience

(3.5 | 3.7 | 3.9)

3.)Interlinking parallel projects on similar topics

(3.2 | 3.5 | 3.8)

4.)Linking up experts within the framework of the project

(3.2 | 3.5 | 3.8)

5.)Finding partners for the initial application of the results

(3.2 | 3.5 | 3.8)

6.)Discussing interim results with all participants

(3.0 | 3.4 | 3.7)

7.)Guiding the IWG to an optimum working model

(3.0 | 3.3 | 3.6)

8.)Communicating industrial trends in the project

(2.9 | 3.2 | 3.5)

9.)Adding further industry partners over the course of the

project (2.9 | 3.2 | 3.4)

10.)Preventing disruptive factors in the project

(2.7 | 3.0 | 3.4)

11.) Organizing and allocating the confirmed project funding

among the partners (2.6 | 3.0 | 3.4)

12.) Procuring test material for the project (2.5 | 3.0 | 3.5)

13.) Monitoring project tasks (2.5 | 2.9 | 3.3)

14.)Maintaining and increasing motivation in the project

(2.4 | 2.8 | 3.2)

15.) Mediating in the event of conflict (2.4 | 2.8 | 3.2)

16.) Supporting the formulation of results at the end of the

project (2.4 | 2.8 | 3.2)

17.) Including marketing organizations linked to the topic

(2.3 | 2.8 | 3.2)

Table 2. How the heads of research institutes view the IWG and the IB role

Research institutes’ expectations of industrial working groups (IWG)

Rank correlation coefficient

Tasks of the innovation broker (IB)

IWGs provide an opportunity to give the research project a practical orientation

0.447* 0.414*

Guiding the IWG to an optimum working modelCommunicating industrial trends in the project

IWGs provide us and the representatives from industry with an opportunity for joint know-how

development on the research topic

0.631** 0.513*

Linking up experts on the topic in the projectInterlinking parallel projects on similar topics

IWGs enable a practice-oriented evaluation of the relevant research progress

0.405* Guiding the IWG to an optimum working model

IWGs provide an opportunity to present the activities of our research institute to industry

0.489* Linking up experts on the topic in the project

Table 3. Rank correlation: what research institutes expect of IWG and IB tasks

* significant at the 0.05 level | ** significant at the 0.01 level


the research institute heads, continueto act as a mediator between the par-ties involved (2.9 | 3.2 | 3.5).

5.2 Industry’s expectations of the IWGand the related IB tasks

In the survey of industry representa-tives, using the four-point Likert scalealready described, the respondentswere asked at the start about the im-portance of the various reasons fortheir participation in an IWG. The re-sults are shown in the left-hand col-umn of Table 4 in descending orderbased on mean values determinedempirically. In addition, as with theresearch institutes, the upper andlower limits of the related 95 % con -fidence intervals (CI) are indicated.The right-hand side of Table 4 shows,on the basis of the same system ofclassification, the results of the as-


sessment of possible IB tasks by theindustry representatives, also with afour-point Likert scale.

The mean values for industry’sexpectations of the IWGs range from1.8 to 3.6, with the related 95 % CIsvarying from 1.6 to 3.7. Except forthe last reason for participation listed,there is a generally favourable assess-ment of the various aspects, althoughit is noticeable that from the 9th as-pect onwards, the IWG expectationsfall short of the mean of 3.0, with thelower limit of the 95 % CI decliningfrom 2.9 to 2.6. The individual assess-ments for aspects 10 to 12 fall corre-spondingly further.

The mean values for the assess-ment of possible IB tasks by the in-dustry representatives questionedrange from 2.6 to 3.6, as shown in theright-hand column of Table 4; thecorresponding 95 % CIs vary from

2.3 to 3.8. Thus, here, too, there is agenerally favourable assessment ofthe possible IB tasks, although an in-crease in the 95 % CI from task 10onwards is apparent owing to thedrop in the lower limit from 2.9 to2.7. The mean of 3.0 is also notreached for the first time from task 10onwards, albeit remaining constant at2.9 up to task 14, with the lower limitof the 95 % CI dropping to 2.7 and2.6 respectively.

In the second evaluation stage,as before, those IWG aspects and IBtasks that had been given a simulta-neous assessment tending towardsthe favourable by the persons polledwere identified. Within the results ob-tained, three areas can be distin-guished: (I) 70 % to 95 % of industryrepresentatives gave a simultaneousassessment tending towards favourablefor both one of the first eight IWG as-


Industry‘s view of the role of the innovation broker (IB) in the project


1.) Publishing the final report (3.4 | 3.6 | 3.8)

2.)Discussing interim results with all participants

(3.2 | 3.4 | 3.6)

3.)Organizing events with the relevant target audience

(3.1 | 3.3 | 3.5)

4.)Linking up experts within the framework of the project

(3.1 | 3.3 | 3.5)

5.)Supporting the formulation of results at the end of the

project (3.0 | 3.2 | 3.4)

6.)Communicating industrial trends in the project

(3.0 | 3.2 | 3.4)

7.)Interlinking parallel projects on similar topics

(3.0 | 3.2 | 3.3)

8.)Finding partners for the initial application of the

results (3.0 | 3.2 | 3.4)

9.)Preventing disruptive factors in the project

(2.9 | 3.1 | 3.3)

10.)Maintaining and increasing motivation in the project

(2.7 | 2.9 | 3.1)

11.) Monitoring project tasks (2.7 | 2.9 | 3.1)

12.) Guiding the IWG to an optimum working model

(2.7 | 2.9 | 3.1)

13.) Adding further industry partners over the course to the

project (2.6 | 2.9 | 3.1)

14.)Organizing and allocating the confirmed project funding

among the partners (2.6 | 2.9 | 3.1)

15.) Mediating in the event of conflict (2.5 | 2.8 | 3.0)

16.) Procuring test material for the project (2.5 | 2.7 | 3.0)

17.) Including marketing organizations linked to the topic

(2.3 | 2.6 | 2.9)

Industry’s reasons for taking part in an industry working group (IWG)


1.)Sharing and increasing technical knowledge

(3.4 | 3.6 | 3.7)

2.)Detecting early trends in technical development

(3.2 | 3.4 | 3.6)

3.)Making contacts with other project participants

(3.1 | 3.2 | 3.4)

4.)Testing new findings for possible implementation

within their own company (3.0 | 3.2 | 3.5)

5.)Obtaining information about the project as a whole on

an ongoing basis (3.0 | 3.2 | 3.5)

6.)Obtaining information on certain individual aspects

of the tasks set (3.0 | 3.2 | 3.5)

7.)Intensifying innovation activities within their own

company (3.0 | 3.2 | 3.4)

8.)Exerting an influence on project progression

(2.9 | 3.1 | 3.2)

9.)Supporting R&D activities in steel applications

generally (2.6 | 2.9 | 3.1)

10.)Helping to shape findings that flow into standardization

(2.6 | 2.8 | 3.1)

11.)Encouraging new research projects on steel application

(2.5 | 2.7 | 3.0)

12.) Obtaining information about competitors (1.6 | 1.8 | 2.0)

Table 4. Industry’s view of the IWG and the IB role



pects and all of the first six IB tasks;(II) 70 % to 85 % of those polled gave asimultaneous assessment tending to-wards favourable for both one of thefirst six IWG aspects and IB tasks 7to 9; (III) 70 % to 75 % of industry rep-resentatives gave a simultaneous as-sessment tending towards favourablefor both one of the first two IWG as-pects and IB tasks 10 to 14. Thismeans that, overall, well over two-thirds of those surveyed who favouredthe first eight IWG aspects at thesame time expressed their approval ofthe first 14 IB tasks with the gradingdescribed. A simultaneously favourableassessment in the range of 12 to 66 %of those questioned was establishedfor all other IWG aspects and IBtasks. For further analysis, in the lightof the broad simultaneous approvalamong industry representatives, thefirst eight IWG expectations are inter-preted as important and the first 14IB tasks as desirable. Here, too, thisinterpretation can also be derivedfrom the mean values and 95 % CIsdescribed above.

The third stage of evaluationalso aims to ascertain the rank corre-lations verifiable by calculation be-tween the aspects interpreted as im-portant and the aspects interpreted asdesirable, and from this to generatetrend-forming assertions regarding apossible connection. Table 5 summa-rizes the aspects identified, showingthe respective rank correlation coeffi-cient and the related significancelevel.

Here, too, based on the coeffi-cients ascertained, which lie between0.30 and 0.49, a moderately strongpositive connection between the rank-ings can be assumed. A direct com-parison with the coefficients in Table 3is not possible, however, because thesample size in the survey of industry

representatives is higher than that ofthe research institute heads.

6 Interpretation: project expectations, IBtask profile, recommendations for CR

6.1 Varying project expectationsamong the groups surveyed

The research institutes’ expectationsof an IWG which are interpreted asimportant can be represented bymeans of three different directions ofinterest for the project phase in CR.Firstly, this includes, in particular,project-related aspects such as theIWG’s role in ensuring that the pro-ject has a practical orientation, thepractice-oriented evaluation of the re-spective research project’s progress inthe IWG and the possible access totopic-related additional informationdisregarded in the project to date. Ad-ditionally, the pursuit of personal in-terests by the respondents can beidentified, which includes the jointdevelopment of new knowledge re-lated to the research topic with theIWG members. Moreover, interestscan be recognized from the IWG ex-pectations which relate to the researchestablishment as an institution. Thisincludes, above all, the benefit of theIWG meetings as milestones and use-ful experience for research staff, thescope for generating new projectsthrough cooperation with the IWGand the opportunity to present the re-search establishment’s project activi-ties and range of work to the IWGmembers (from industry).

The reasons for industry’s par-ticipation interpreted as important inan IWG accompanying a CR projectcan also be summarized based on thethree different directions of interestdescribed above. Whereas a relativelyuniform weighting within the IWGaspects interpreted as important is

discernible among the research insti-tutes, the industry representatives dis-play a clear prioritization of personalinterests in their IWG participationin the form of sharing and increasingtheir own technical knowledge of theresearch topic. Following close be-hind, company-related interests areimportant to the respondents. This in-cludes, in particular, detecting newtechnical trends, keeping in contactwith other project participants alsointerested in the topic, monitoringthe results for possible application intheir company and generally intensi-fying their company’s innovation ac-tivities. In third place among the IWGaspects deemed important by the in-dustry representatives is their project-related interests. This includes, inparticular, receiving general projectinformation on an ongoing basis, butalso being able to pursue very specificaspects of the research topic in theIWG. Additionally, the purpose of ex-erting influence on project progres-sion through IWG participation,identified as important, should bementioned in this context. Industryrepresentatives regard obtaining in-formation about potential competi-tors within the IWGs as somewhatunimportant. Fig. 2 provides anothersummary of the varying interests ofscience and industry with regard totheir IWG participation.

6.2 Developing the IB task profile forsteel application research in theconstruction industry

On the basis of the IB tasks inter-preted as desirable, the following IBtask profile for pre-competitive CRon steel applications in the construc-tion sector can be drawn up based onthe various key person roles outlinedat the start:


Industry’s reasons for taking part in an industrial working group (IWG)

Rank correlation coefficient

Tasks of the innovation broker (IB)

Sharing and increasing technical knowledge 0.305* Organizing events with the relevant target audience

Detecting early trends in technical development 0.334* Interlinking parallel projects on similar topics

Making contacts with other project participants0.398**0.327*

Interlinking parallel projects on similar topics Linking up experts on the topic in the project

Obtaining information about the project as a whole on an ongoing basis

0.316* 0.314*

Interlinking parallel projects on similar topicsDiscussing interim results with all participants

Exerting an influence on project progression 0.312* Monitoring the project tasks

Table 5. Rank correlation: industry’s reasons for taking part in the IWG and IB tasks

* significant at the 0.05 level | ** significant at the 0.01 level


In the early innovation phases,the IB’s tasks consist chiefly of activeideas management in which the IBinitiates or accepts and collects ideaoutlines, taking them to the point ofindustrial pre-evaluation of an initialproposal. The IB also tries very earlyon, within the scope of network man-agement, to arrange industry contactsto support the idea and promotes theforging of collaboration between theresearch institutes and the relevantindustry representatives. At the verystart of the innovation process, the IBis thus performing the groundworkaimed at preventing the possible weak-ness – outlined at the start – of a lackof practical orientation in the project.This activity, not yet completely pro-ject-related, can be likened to that ofa (product) champion role at the con-ception stage, which is characteristicof early support for an innovationand through which an idea is pro-moted based on conviction and withthe political skill often required. Aspart of early ideas management, theIB also tries to initiate and maintaincommunication between the partiesfrom science and industry, thus help-ing to define clearly the project goalsof both players. Possibly varying ex-pectations of a project criticized atthe outset are thus harmonized earlyon and a clear starting point for the


desired project phase comes about.All in all, the activities of the IB dur-ing the early innovation phases canbe summed up as those of a coordi-nating and contact point for innova-tions. Haller [34] appropriately de-scribes the challenge of such a posi-tion in “concentrating on dealingswith ideas generators and in support-ing their individual needs related tothe finding, formulation and imple-mentation of new innovative ideas”.In return, the IB confirms a lasting in-crease in the quality of new researchinitiatives.

The IB engages in the subse-quent arranging of talks with possiblefunding providers as part of activefunding management, which basedon the results can be interpreted asthe most important aspect of the ad-ministrative innovation phases fromthe research institutes’ perspective.For this activity, the IB must recog-nize the process steps required in theadministrative phases of a pre-com-petitive CR project and play a sup-portive role, taking account of the of-ten differing requirements of the vari-ous funding providers in the processof seeking to attract funds. The IB’stasks in this instance can be likenedto that of a process promoter, who inthis situation is acting as a translatorbetween the linguistic worlds of tech-

nology and economics. On the basisof possible experience from previousCR projects, the IB can also use theknowledge of specific funding pro -viders across projects, which will givethe actual project-related promoterrole impetus over time.

In the project phase, the IB’s ac-tivities are related to the project levelon the one hand but also requirecross-project action on the other. Onthe basis of the research institutes’view, the IB’s tasks can be assigned tofive different management areas. Forthe project phase, as indeed alreadydescribed for the early innovationphases, this means that the IB shouldengage in active network manage-ment for the project. The research in-stitutes welcome support with findingand linking-up suitable experts fromthe relevant specialist fields, who arein turn willing to supervise a specificproject. In addition, the IB should inthis context assist in identifying com-panies that may be interested in im-plementing the research results andshould promote the recruitment ofadditional participants for the IWGduring the course of the project. An-other element in the IB’s tasks duringthe project phase is topic-related in-formation management, which helpsthe research institutes to obtain infor-mation about parallel projects onsimilar topics and find out about in-dustry trends in the course of the pro-ject. Also important to the researchinstitutes is support for discussing in-terim results, which can be achievedthrough active knowledge manage-ment by the IB. Correspondinglyadapted team management, whichhelps the IB to guide the projectgroup, including the IWG, to an opti-mum working model and in whichdisruptive factors in the group areprevented early on or jointly resolved,are also regarded as desirable by theresearch institutes.

There is a degree of consensusbetween the industry representatives’views and the research institutes’ as-sessments in relation to the projectphase. This concerns the IB tasks ofnetwork management and informa-tion management. In addition, the in-dustry representatives regard it as de-sirable for the IB – as part of the IB’sknowledge management activities andapart from supporting discussions oninterim results – to assist with the for-


Fig. 2. Varying interests of science and industry in an IWG accompanying a CRproject



mulation of the research results at theend of the project as well. With theIB’s team management activities, theindustry representatives also see theneed for the IB, in addition to steer-ing the IWG and preventing disrup-tive factors within it, to play an activerole in maintaining and increasingmotivation among the project part-ners during the project. As an addi-tional element in the IB’s tasks, theindustry representatives welcome theassumption of responsibility for pro-ject management aspects, such as themonitoring of project tasks alreadyagreed among the participants andthe performance-related payment offunds approved at the start of the pro-ject over the project term.

Based on the IB tasks deter-mined from the areas of knowledgemanagement and information man-agement, some of the IB’s activitiesfor the project phase can be summa-rized as those of a know-how pro-moter. Team management and projectmanagement tasks, which are amongthe tasks performed by a process pro-moter, were also identified as desir-able for the project phase. The addi-tional, desired use of active networkmanagement by the IB links the IBrole to that of a relationship pro-moter.

Summing up, it can be seen thatthe IB, even during the project phase,in addition to supporting the project’spractical orientation, makes a sub-stantial contribution to the know-howsought by science and industry and tothe exchange of information amongthe project participants. The IB thusmakes a significant contribution toeliminating the disadvantages causedby the “different languages” spokenby science and industry, as criticizedat the outset, and to clarifying andharmonizing the differing expecta-tions of a CR project. The IB’s role inthe project phase can thus be inter-preted as that of an intermediate per-son between research institutes andindustry project partners. This meetsthe requirements of the role requiredby Abbott et al. [16] for the construc-tion sector of a person who managesinteraction at the university-businessinterface. At the same time, the IB’sactivities in the IWG support the ef-fectiveness when transferring results.Blayse et al. [33] describe how impor-tant it is for the achievement of a suc-

cessful transfer in the constructionsector that “the industry partners un-derstand the R&D results and arethen actually put in a position to usethese results (in day-to-day construc-tion business)”.

For the late innovation phases,the central publication and dissemi-nation of the final report, as well asthe holding of events on the researchtopic, were stressed as desirable IBtasks by both the research institutesand the industry representatives. Aspart of active dissemination manage-ment, the IB thus supports the re-search institutes in disseminating theirresearch results and creates – beyondthe duration of the project – addi-tional access to the expertise created.For this purpose, the IB should be ableto demonstrate specialized knowledgeof the results and the IB’s role in thelate innovation phases could there-fore be likened to that of a know-how promoter. The additional tasksmentioned regarding the organizationof a target group-oriented event andfinding suitable partners for the ini-tial application of the research find-ings can be assigned to the role of arelationship promoter.

If the CR phase model shown inFig. 1 is examined against the back-ground of the IB task profile de-scribed, the varying roles of the IB ina comprehensive examination of theinnovation process reveal that the IBnot only operates at the micro-level,i.e. in the phases themselves, but alsofrequently ensures the innovation-re-lated flow of information beyond thephase boundaries at the macro-level,thus assuming a prominent role in in-terface management between the in-dividual innovation phases and thedifferent CR participants. His activi-ties are therefore also similar acrossall phases to that of a gatekeeperwho is endeavouring to push an inno-vation process forward from the ideato the application of the findings andto supplement this process with addi-tional topic-related information.

Fig. 3 summarizes the IB tasksidentified as desirable by science andindustry for CR on steel applicationsin the construction industry. In addi-tion, the IB’s changing roles throughthe individual innovation phases areshown.

The comprehensive IB task pro-file depicted shows an initial approach

to the implementation of the changein the results transfer in applied re-search in the construction sector, asdemanded by Abbott et al. [16]. Theydescribe this as follows: “there is aneed to move away from the old fash-ioned notion of the university devel-oping new ideas through research,and then transferring and disseminat-ing these ideas to (construction) in-dustry in a linear fashion”.

6.3 General trends in CR management

The IWG expectations assessed asimportant (see Fig. 2) indicate thevarying interests of the project parti -cipants from science and industry. Inthe same way, the IB tasks assessedas desirable by the respondents (seeFig. 3) form the basis for creating anarea of activity for the innovationbroker in CR on steel applications inthe construction industry. In additionto this opinion, on the basis of the es-tablished (highly) significant correla-tions between the IWG expectationsassessed as important and the IB tasksassessed as desirable in accordancewith Tables 3 and 5, the followingtrend-forming assertions can be madeabout CR management:

Based on the results of the sur-vey of heads of research institutes,one can assert the following trendthat those research institutes that …

… regard a practice-oriented ap-proach to the project by the IWG asimportant, to a significant degree alsoregard it as desirable that the IBshould guide the IWG to an optimumworking model in the project andcommunicate trends from the indus-try involved in the project.

… regard the joint developmentof know-how with industry in theIWG as important, to a highly signifi-cant degree also regard it as desirablethat the IB should link up experts onthe research topic and (to a signifi-cant degree) interlink parallel pro-jects on similar topics.

… regard a practice-oriented eval-uation of the interim results withinthe IWG as important, to a signifi-cant degree also regard it as desirablethat the IB should assess the guidingof the IWG to an optimum workingmodel.

… regard the possibility of pre-senting their activities in the IWG asimportant, to a significant degree also



regard it as desirable that the IB shouldassess a link-up of experts on the re-search topic.

Based on the results of the industrysurvey, one can assert the followingtrend that those industry representa-tives who …

… regard the exchange of tech-nical knowledge in the IWG as impor-tant, to a significant degree also regardit as desirable that the IB should as-sess the holding of an event during orafter the project term with the righttarget audience.

… regard the detection of techni-cal trends through participation inthe IWG as important, to a signifi-cant degree also regard it as desirablethat the IB should assess an interlink-ing of parallel projects on similar top-ics.

… regard the contacts establishedwith other project participants withinthe IWG as important, to a highly sig-nificant degree also regard it as desir-able that the IB should assess a link-up of parallel projects on similar top-ics and (to a significant degree) alink-up of experts in the project topic.


… regard the continuous provi-sion of information on the researchproject in the IWG as important, to asignificant degree also regard it as desir-able that the IB should assess the inter-linking of parallel research projects onsimilar topics and supporting the dis-cussion of interim results in the IWG.

… regard the possibility of exert-ing influence on the project throughtheir participation in the IWG as im-portant, to a significant degree also re-gard it as desirable that the IB shouldassess the allocation and monitoringof tasks in the IWG.

7 Conclusions

Innovation requires not only the gen-eration of an idea, but also its suc-cessful implementation [12]. This mustalso apply, in particular, to pre-com-petitive CR between the steel indus-try and the construction sector, theresults of which will benefit a largenumber of companies. In order toachieve this goal, it is necessary toovercome the deficiencies in the ex-isting CR system mentioned at thestart and also to take into account

the special features of the construc-tion sector.

In the light of this, the steel in-dustry in Germany, as a key supplierof materials and innovations for theconstruction sector, is seeking a com-prehensive optimization of CR man-agement for steel applications in theconstruction industry. This article sug-gests an approach that consists largelyof using a central IB role within thewhole CR innovation process.

The various project expectationsof the two groups were identifiedthrough a survey of representativesfrom the scientific community and in-dustry in a FOSTA research clusterwhich accompanied the project. Insummary, these expectations are dri-ven by personal, organization-relatedand project-related interests.

On the basis of the survey re-sults, a task profile has been devel-oped for the IB which relates to thewhole CR innovation process – fromfinding ideas to implementing the re-sults obtained. The IB activities inter-preted as desirable come from the areas of idea management, fundingmanagement, network management,


Fig. 3. The tasks and changing roles of the IB in the CR innovation process



information management, team man-agement, project management, know -ledge management and disseminationmanagement, and are distributedacross the whole process at both pro-ject and cross-project level.

The IB task profile created al-lows an assignment to familiar keyperson roles according to the con-tent of the individual subtasks. It hasemerged, however, that the IB for CRon steel applications in the construc-tion industry cannot be assigned fullyto any of the existing key personroles. Rather, the IB’s area of activityrepresents a mixture of existing andnew elements. Instead of a rather sta-tic assignment of a role to a field ofactivity, in this case a dynamic cross -over is discernible for CR betweenthe varying roles of the different keyperson concepts during the course ofthe innovation process. On the basisof this dynamic component and theassociated increase in the interfaceproblems between the individual ele-ments of the IB task profile, depart-ing from the one-person concept forthe key person in CR on steel appli-cations in the construction industry isnot expedient.

Finally, trend-forming assertionsabout CR management were drawnup in this study, which arose basedon significant correlations betweenIWG expectations and IB tasks. Theseshow the IB’s firm integration intoCR and the acceptance of the IB’s activities throughout the innovationprocess.

As part of its further efforts tooptimize the management of the steelindustry’s pre-competitive CR with itscustomers, FOSTA intends to con-tinue implementing and optimizingthe IB task profile developed thus farwithin the NASTA research cluster.In particular, the assumption of thetasks identified for the project phasewould make a significant contribu-tion to achieving success with inno-vations. An evaluation of the profiledeveloped to date through individualpost-project reviews is planned forthe final phase of the six NASTA sub-projects. Within the scope of cross-topic management of steel applica-tion research by FOSTA, there areplans to enable an institutionalizedapplication of the IB concept for fu-ture CR clusters with other steel ap-plication sectors.

Acknowledgements

The management aspects put forwardhere resulted from an ongoing FOSTAresearch project on innovation man-agement in steel application researchwith the participation of the Facultyof Science, Engineering & Comput-ing, Kingston University London, andthe Faculty of Management, SocialWork & Construction of the Univer-sity of Applied Sciences & Arts, Holz-minden. FOSTA is grateful to the re-search institutes and industrial part-ners of the NASTA project for theirparticipation in the survey. The NASTAproject itself is supported by the Fed-eral Ministry of Economics & Tech-nology through the AiF as part of theprogramme for promoting industrialcooperative research (IGF) in accor-dance with a resolution of the lowerhouse of the German parliament.

References

[1] WR (2007) Empfehlungen zur Inter-aktion von Wissenschaft und Wirtschaft.Deutscher Wissenschaftsrat, Oldenburg.

[2] RWI (2009) Erweiterte Erfolgskon-trolle beim Programm zur Förderungder IGF im Zeitraum 2005–2009. RWIEssen on behalf of the Federal Min-istry of Economics & Technology.

[3] Nuesse, G., Limbachiya, M., Herr, R.,Wieland, H. J.: Active management ofthe early innovation phases in steel ap-plication research for the constructionsector. Steel Construction – Design andResearch, vol. 4, No. 1, 2011, pp. 34–40.

[4] Nuesse, G., Limbachiya, M., Herr, R.:Integrated management of steel appli-cation research – SME appropriate co-operation between the steel industryand the construction sector. wissen -schaftsmanagement, vol. 16, No. 1,2010, pp. 43–49.

[5] Reichstein, T., Salter, A., Gann, D.:Break on Through: Sources and Deter-minants of Product and Process Inno-vation among UK Construction Firms.Industry and Innovation. vol. 15, No. 6,2008, pp. 601–625.

[6] Fairclough, J.: Rethinking construc-tion innovation and research. Depart-ment of Trade & Industry, London,2002.

[7] Barrett, P., Sexton, M., Lee, A.: Inno-vation in Small Construction Firms.Taylor & Francis, London, 2008.

[8] Slaughter, S.: Models of ConstructionInnovation. Journal of ConstructionEngineering and Management. IssueMay/June 1998, pp. 226–231.

[9] Harris, M., Halkett, R.: Hidden Inno-vation – How innovation happens insix “low innovation” sectors.: NESTA –National Endowment for Science, Tech-nology and the Arts, London, 2007.

[10] Reichstein, T., Salter, A., Gann, D.:Last among equals: a comparison of innovation in construction, services andmanufacturing in the UK. ConstructionManagement and Economics. vol. 23,No. 6, 2005, pp. 631–644.

[11] Sexton, M., Barrett, P., Aouad, G.:Motivating small construction compa-nies to adopt new technology. BuildingResearch & Information. vol. 34, No. 1,2006, pp. 11–22.

[12] Sexton, M., Barrett, P.: Appropriateinnovation in small construction firms.Construction Management and Econo -mics. vol. 21, No. 6, 2003, pp. 623–633.

[13] emcc: Trends and drivers of changein the European construction sector:Mapping report. Dublin: EuropeanMonitoring Centre on Change (emcc) &European Foundation for the Improve-ment of Living and Working Condi-tions, 2005.

[14] Haller, M., Lawson, M., Nuesse, G.,Tengberg, C., Beguin, P., Ungermann,D., Kalameya, J.: EuroBuild in Steel –Evaluation of client demand, sustain-ability and future regulations on thenext generation of building design. European Commission, RFCS ProjectNo. 7210-PR/381, Brussels, 2007.

[15] Barnes, T., Pashby, I., Gibbons, A.:Effective University–Industry Interac-tion: A Multi-case Evaluation of Col-laborative R&D Projects. EuropeanManagement Journal, vol. 20, No. 3,2002, pp. 272–285.

[16] Abbott, C., Aouad, G., Madubuko,L.: An Innovation Platform for Con-struction. Manchester: North WestUniversities Association (nwua), 2008.

[17] Couchman, K., Beckett, R.: Achiev-ing Effective Cross-Sector R&D Col-laboration: A Proposed ManagementFramework. Prometheus, vol. 24, No. 2,2006, pp. 151–168.

[18] Nuesse, G., Wieland, H. J., Feld-mann, M.: Nachhaltiges Bauen, Stahl -bau, vol. 80, No. 10, 2011, pp. 701–702.

[19] Schon, D. A.: Champions for Radi-cal New Inventions. Harvard BusinessReview. vol. 41, No. 2, 1963, pp. 77–86.

[20] Chakrabarti, A. K.: The Role ofChampion in Product Innovation. Cal-ifornia Management Review, vol. 17,No. 2, 1974, pp. 58–63.

[21] Shane, S. A.: Are Champions Differ-ent from Non-Champions? Journal ofBusiness Venturing, vol. 9, No. 5, 1994,pp. 397–421.

[22] Markham, S. K.: Corporate Cham-pioning and Antagonism as Forms of



Political Behaviour: An R&D Pers -pective. Organization Science, vol. 11,No. 4, 2000, pp. 429–447.

[23] Witte, E.: Organisation für Innova-tionsentscheidungen – Das Promoto -ren-Modell. Otto Schwarz, Göttingen,1973.

[24] Hauschildt, J., Chakrabarti, A. K.:Arbeitsteilung im Innovationsmanage-ment – Forschungsergebnisse – Krite-rien und Modelle, Zeitschrift Führungund Organisation, vol. 57, No. 6, 1988,pp. 378–389.

[25] Gemünden, H. G., Walter, A.: För -derung des Technologietransfers durchBeziehungspromotoren. Zeitschrift Füh -rung und Organisation, vol. 65, No. 4,1996, pp. 237–245.

[26] Walter, A.: Der Beziehungspromo-tor. Ein personaler Gestaltungsansatzfür das Relationship Marketing. Gabler,Wiesbaden, 1998.

[27] Folkerts, L.: Promotoren in Inno -vationsprozessen: Empirische Untersu -chung zur personellen Dynamik, Dis-sertation. Deutscher Universitäts-Ver-lag, Kiel, 2001.

[28] Allen, T. J.: Communication net-works in R&D laboratories, R&D Man-agement, vol. 1, No. 1, 1970, pp. 14–21.


[29] Domsch, M., Gerpott, H., Gerpott, T.:Technologische Gatekeeper in der in-dustrielle F&E – Merkmale und Leis-tungswirkungen. Poeschel, Stuttgart,1989.

[30] Hauschildt, J., Schewe, G.: Gate-keeper und Prozeßpromotoren. In:Hauschildt, J., Gemünden, H. G.: Pro-motoren – Champions der Innovation.Gabler, Wiesbaden, 1999, pp. 161–176.

[31] Hauschildt, J.: Zur Weiterentwick-lung des Promotoren-Modells. In: Hau -schildt, J., Gemünden, H. G.: Promo-toren – Champions der Innovation.Gabler, Wiesbaden, 1999, pp. 257–283.

[32] Gemünden, H. G., Hölzle, K.: Schlüs-selpersonen der Innovation – Cham -pions und Promotoren. In: Gassmann,O., Albers, S.: Handbuch Technologie-und Innovationsmanagement. Gabler,Wiesbaden, 2005, pp. 459–474.

[33] Blayse, A., Manley, K.: Key influ-ences on construction innovation. Con-struction Innovation, vol. 4, No. 3, 2004,pp. 143–154.

[34] Haller, C.: VerhaltenstheoretischerAnsatz für ein Management von Inno-vationsprozessen, Dissertation. Betriebs -wirtschaftliches Institut der UniversitätStuttgart, 2003.

Keywords: pre-competitive cooperativeresearch; innovation broker; manage-ment of interdisciplinary research; sus-tainability in steel construction; pro-moter

Authors:Dipl.-Ing. Gregor Nuesse MSc Research Association for Steel ApplicationSohnstr. 65, 40237 Düsseldorf, [email protected]

Prof. Mukesh Limbachiya BE MSc PhD CEng FIMS MICE MIMMMFaculty of Science, Engineering and ComputingKingston University LondonPenrhyn Road, Kingston upon Thames,KT1 2EE, UK

Prof. Dr.-Ing. Roland HerrFaculty of Management, Social Work & ConstructionHAWK University of Applied Sciences & Arts Haarmannplatz 3, 37603 Holzminden, Germany

Alan Ellis MSc MCIOB MBEngFaculty of Science, Engineering and ComputingKingston University LondonPenrhyn Road, Kingston upon Thames, KT1 2EE, UK


High quality plate supplied by DillingerHütte for the structural engineering sectoris used in demanding, out of the ordinaryprojects around the world: in the Viaducde Millau in the South of France, in theOlympic stadium in Athens, in the multi-storey car park for Stuttgart’s new exhibi-tion centre, in the „Simone de Beauvoir“footbridge, in Paris, and in Shanghai’sWorld Financial Center – many ofwhich only became possible thanks totailor-made solutions from Dillingen.

Please visit www.dillinger.de for moreinformation.

10th International Structural Steel Colloquium

Around 200 experts in structural steelengineering from around the world res-ponded to Dillinger Hütte’s invitation to the 10th Dillingen Structural SteelColloquium, held in the civic hall in Dillingen on 1 and 2 December 2011.

The attending specialists from struc-tural steel engineering companies, engi-neering consultancies, government de-partments and agencies, universities andresearch institutes discussed current top-ics in the practical use of steel, and thelatest research results in this field. Note-worthy projects implemented using heavyplate from Dillingen were naturally againhigh lighted, among them the MakkahClock Tower, a skyscraper currently un-der construction in Saudi Arabia. Thecentre of interest here is the steel usedand the challenges encountered in suchremark able civil engineering projects. Thesustainable use of steel in civil engineer-ing was, indeed, a particular focus of thiscolloquium.

„Steel is a sustainable material; it canbe recycled 100 % – again and again andagain. As well as this, the use of steel incivil engineering shortens completiontimes considerably, and therefore reduces

Conferencesenvironmental problems, such as noise,dirt and disruptions to traffic, to the veryminimum during bridge construction pro-jects, for instance“, noted Dr. KarlheinzBlessing, Chief Executive Officer ofDillinger Hütte, in his speech of welcome,and, as he continued: „The expansion ofregenerable energy sources would be in-conceivable without innovative steelproducts. This is why it is so importantnot to compromise the competitivenessof the European steel industry with emis-sions rights trading and with the burdensof high energy prices“.



Reports

– northern channel main bridge plusthe northern approach bridge

– southern channel main bridge,southern approach bridge and middle approach bridge.

A bidder was allowed to tender forone section only, rather than bothsections. But each tender had to in-clude the overall general layout ofboth the southern and northern chan-nel bridges due to their close proxim-ity.

A suspension bridge with a mainspan exceeding 900 m was preferredfor the main bridge over the southernchannel.

2.2 Key boundary conditions for main bridge

1) High-voltage corridors on Guang -zhou side of northern channelA high-voltage corridor belong-ing to Huangpu power station islocated on the Guang zhou sideof the northern channel, whichis fixed on plan. The elevationof the high-voltage cables isabout 60 m based on the origi-nal design and can be raised nohigher than 100 m. The distancebetween the high-voltage corri-dor and the Guangzhou bank ofthe northern channel is about90 m.

2) Commercial wharf on Guangzhoubank of northern channelThere is a commercial wharf onthe Guangzhou bank of the north-ern channel which is owned by anoverseas company. The affect onthe commercial wharf must be re-duced to a minimum. The distancebetween pier and bank should begreater than 10 m.

1 Site topography, design standardsand criteria

1.1 Site topography

The Huangpu Pearl River Bridge is thekey project on the second ring high-way at Guangzhou, China. The bridgeconnects the Central Business District(CBD) of Guangzhou and Panyu dis-trict and crosses the Pearl River be-tween the Guangzhou Overseas ShipMaintenance Plant and the BoluomiaoShip Factory. An island about 600 mwide divides Pearl River into a southernand a northern channel at this point.

The total width of the northernchannel is about 390 m and the waterdepth is about 5–8 m. The total widthof the southern channel is about1100 m, the water is 3–30 m deep.

1.2 Design criteria

The key design criteria adopted forthe bridge include: (1) Highway class: 6-lane highway in

the short-term, 8-lane highway inthe long-term.

(2) Bridge width: 34.5 m (excludingcable anchorage zone width).

(3) Design speed: 100 km/h.(4) Imposed load standard: Chinese

highway class I.(5) Navigation clearance: the width

and height of the northern navi-gation channel should not be lessthan 280 and 55 m respectively.The maximum water level for nav-igation is 7.0 m, the minimum wa-

ter level for navigation is 3.48 m.According to the requirements ofthe Guangdong Province Hydrau -lic Department, the minimum spanshould be greater than 380 m.

(6) Design wind speed: average max-imum 10-minute wind speed is38.4 m/s, with a 100-year returnperiod and measured 20 m abovesea level.

(7) According to the site earthquakesafety assessment report, com-pleted by the Guangdong ProvinceSeismic Department, the groundsurface horizontal acceleration is0.228g, with a 100-year return pe-riod and a probability of 3 %.

1.3 Design standards

In accordance with the client’s require-ments, the preliminary and detaileddesign of the permanent works wascarried out according to the Chinesedesign code for highway bridges, theEuropean Union code (Eurocode) andBritish Standard BS 5400 [1] for theentire bridge; other compatible codesand technical literature were adoptedfor specific design issues not coveredby the main standards. The entirebridge was designed according to ulti-mate and serviceability limit states.

2 Tender design for main bridge 2.1 Background to tender design

The bridge was divided into two sec-tions for tendering:

Huangpu Pearl River Northern Channel Bridge, Guangzhou – the longest single-pylon cable-stayedbridge span in ChinaYingliang Wang Chengzao Huang Yuncheng Feng

DOI: 10.1002/stco.201200007

The Huangpu Pearl River Bridge is the key project on the second ring highway of Guang -zhou, China. The main bridge over the northern channel is a single-pylon cable-stayedbridge with a main span of 383 m and a steel box girder, which was constructed usingthe balanced cantilever method. This paper describes the tender design, the final designand the construction of the bridge, including foundation, pylon and main girder.

12_053-060_Wang (SC1)_000-000_SC_Bericht (1).qxd 10.02.12 08:04 Seite 53

54

Reports


3) LandscapingThe landscaping should be consid-ered as a unity due to the close re-lationship between the NorthernChannel Bridge and the SouthernChannel Bridge

2.3 Tender design proposals

A total of 24 proposals were submittedby seven Chinese tenderers, which areshown in Table 1. Fig. 1 shows thegeneral layout of proposals 1 to 5.

2.4 Tender design results

A proposal for the northern channelusing a double-pylon cable-stayedbridge with a span of 450 m was ex-cluded because the side span cableon the Guangzhou side would inter-sect with the high-voltage cables. Asuspension bridge proposal was alsorejected: although its cables did notclash with the high-voltage cables, theanchorage on the Guangzhou sidewould have affected the wharf. In the

end, a single-pylon cable-stayed bridgewith the pylon located on the islandwas approved by the expert and theclient.

The design of the Southern Chan-nel Bridge, southern and middle ap-proach bridges was awarded to CCCCHighway Consultants Co. Ltd. The de-sign of the Northern Channel Bridgeand the northern approach bridgeswas awarded to CCCC First HighwayConsultants Co. Ltd. The entire inde-pendent checking of the design and

Pro-posal Proposals for Northern Channel BridgeProposals for Southern

Channel BridgeConsultant

1 SPCSB with 398 m main span and steel box girder SB with 1100 or 900 m main span

Joint venture of Highway Planning &

Design Institute, Sichuan province,

and Chongqing Communications

Research & Design Institute Co. Ltd.

2SPCSB with 398 m main span and

steel-concrete hybrid girderSB with 1100 or 900 m main span

3 SB with 468 m main span and steel box girder SB with 1100 m main span

4SPCSB with 398 m main span and steel box girder,

connected to the suspension anchorageSB with 1100 m main span

5DPCSB with 450 m main span and

steel-concrete hybrid girderSB with 1100 m main span


concrete box girderSB with 1100 m main span



CCCC First Highway Consultants Co. Ltd.

8 DPCSB with 456 m main span and steel box girder SB with 900 m main span


steel-concrete hybrid girderSB with 900 m main span



Highway Planning & Design Institute,

Guangdong province


steel concrete hybrid girderSB with 900 m main span




14 DPCSB with 450 m main span and steel box girder SB with 900 m main span Highway Consultants, Anhui province15 DPCSB with 360 m main span and steel box girder SB with 900 m main span


CCCC Highway Consultants Co. Ltd.

17 DPCSB with 450 m main span and steel box girder SB with 760 + 760 m main span**

18 DPCSB with 450 m main span and steel box girder SB with 960 + 380 m main span

19 DPCSB with 450 m main span and steel box girder DPCSB with 890 m main span

20 DPCSB with 450 m main span and steel box girder SB with 1096 m main span China Railway Major Bridge

Reconnaissance & Design

Institute Co. Ltd.



23 DPCSB with 450 m main span and steel box girder DPCSB with 900 m main span

24 N/A SB with 900 m main span SMEDI

Notes DPCSB: double-pylon cable-stayed bridge; SPCSB: single-pylon cable-stayed bridge; SB: suspension bridge;

SMEDI: Shanghai Municipal Engineering Design Institute; ** triple-pylon suspension bridge

Table 1. Proposals for Pearl River Bridge



Reports

construction stages was awarded tothe joint venture of Highway Plan-ning & Design Institute, the Depart-ment of Communication, Sichuanprovince, and Chongqing Communi-

cations Research & Design InstituteCo. Ltd. The first author was ap-pointed as chief independent check-ing consultant by the client for hiscontribution to proposals 1, 2 and 3.

3 Final proposal and landscaping3.1 Final design of complete crossing

At the final design stage, the mainspan of the northern and southern

Fig. 1. General layout of the proposals at tender stage (unit: cm)

North channel bridge South channel bridge


56

Reports


channel bridges were adjusted to 383and 1108 m respectively. Fig. 2 showsthe general layout of the entire cross-ing.

The total length of the HuangpuPearl River Bridge is 7049 m. Thebridge is divided into:1) Northern approach bridge – a pre-

stressed concrete continuous girderwith a span arrangement of 11 ×30 + (30 + 5 × 45) + 6 × 45 + (45 +2 × 62.5) + 2 × 62.5 + 5 × 62.5 +5 × 62.5 m

2) Northern channel main bridge – asteel box girder cable-stayed bridgewith a span arrangement of 383 +322 m

3) Middle approach bridge – a pre-stressed concrete frame with a spanarrangement of 6 × 62.5 m

4) Southern channel main bridge – asteel box girder single-span sus-pension bridge with a main spanof 1108 m

5) Southern channel approach bridge– a prestressed concrete continu-ous girder with a span arrange-ment of 6 × 62.5 + 5 × 62.5 + 5 ×62.5 + (4 × 62.5 + 45) + 6 × 45 +

6 × 45 + (5 × 45 + 30) + 7 × 30 +7 × 30 + 7 × 30 + 6 × 30 + 6 ×30 m.

3.2 Landscaping for main bridge

The landscaping had to consider thebridge as a whole and keep the con-figuration of the pylons harmonizedfor both the northern and southernchannel main bridges. Although abidder was allowed to tender for onesection only, rather than both sec-tions with their different bridge types,they are connected by the middle ap-proach bridge with a length of just375 m. Therefore, a similar pylon con-figuration was adopted. It is very sim-ilar to the Chinese character “MEN”,which means the “gate of Guangzhou”.Fig. 3 shows a photomontage of themain bridge.

4 Design of northern channel main bridge

4.1 General design

The main bridge for the northernchannel is a single-pylon cable-stayed

bridge with a steel box girder and aspan arrangement of 383 + 197 + 62.5+ 62.5 m (Fig. 4). The ratio of sidespan length to centre span length is0.8407. Two intermediate piers wereadopted to enhance the general rigid-ity and improve the structural behav-iour of main girder and pylon. Theapproach bridge uses prestressed con-tinuous box girders with main spansof 62.5 and 45 m, which are not in-cluded in this paper.

The pylon is located in the shal-low water zone, pier 38 is located onthe riverbank to avoid affecting thecommercial wharf, intermediate piers40 and 41 are located on the middleisland. In order to avoid uplift of theside span, a counterweight is includedin this span.

Vertical bearings are included atthe pylon and piers 39, 40 and 41;there is a traverse wind bearing onthe inner side of the pylon shaft andlongitudinal hydraulic dampers at thepylon cross-beam (Fig. 5).

4.2 Main girder

The main girder is a streamlined steelbox girder with three cells (Fig. 6).The depth of the box girder is 3.5 mand the total width 41 m, which in-cludes the windshield and the cableanchorage zone. The transverse fall is2 %. The top and bottom plate thick-nesses of standard segments are 16and 12 mm respectively, and both ofthem are 20 mm thick near the pylonsegment.

Double inner webs divide thebox into three cells; the distance be-tween inner web and centre line ofbridge is 9.4 m. The longitudinal stiff-eners are closed trapezoidal stiffeners.The height, centre-to-centre distanceFig. 3. Photomontage of main bridge

Fig. 2. Elevation and plan of entire bridge, including the locations of piers, abutment and approach bridge (unit: m)



Reports

and thickness of the top flange longitu-dinal stiffener are 280, 600 and 8 mmrespectively. The height, centre-to-cen-tre distance and thickness of the bot-tom flange longitudinal stiffener are250, 750 and 6 mm respectively.

The spacing of the cross-beams is3.2 m, their web thickness is 10 mm,but 14 and 22 mm at the cable an-chorages and bearings respectively.The cross-beam is divided into threesegments in the transverse direction.

The segment length of the steelbox girder was governed by the capac-ity of the floating crane, the derrick, aswell as the construction method andduration of the work. The steel boxgirder is divided into 52 segments;the standard segment length is 16 m,the maximum weight 321 t.

The site splices of the steel boxgirder are welded. The total weight ofthe steel box girder is 15 500 t. GradeQ345qC has been adopted for the steelbox girder, which is required to havea yield strength of 345 MPa and animpact test energy of 41 J at 0 °C.

The steel orthotropic deck sys-tem was analysed using the Pelikan-Esslinger method; the analysis modelis shown in Fig. 7. Table 2 lists thestresses in the deck system.

4.3 Anchorage of stay-cables at main girder

Figs. 8 and 9 show the anchorage de-vice for the stay-cable at the maingirder.

The anchorage device was weldedto the external side of the web of thesteel box girder using full penetrationwelds. The plate of the external webof the steel box girder requires a Z25through-thickness property to avoidlamellar tearing. The through-thick-ness property of the plate was deter-mined with Pre-EN 1993-1-10.

The local stresses in the anchor-age region were analysed using the fi-nite element method (Fig. 10).

Fatigue assessment of the full pen-etration welds under tensile and shearstress interaction was checked usingEN 1993-1-9 and the French steelbridge fatigue code (Ponts métalliqueset mixtes: Résistance à la fatigue).

4.4 Pylon and foundation

The configuration of the pylon is sim-ilar to the pylon of the southern chan-

Fig. 4. General layout of northern channel main bridge (unit: cm)

Fig. 5. Bearing system: a) general layout of bridge, b) bearings at pylon cross-beam(unit: mm)

Fig. 6. Section through standard segment steel box girder (unit: mm)

Location

Stress

Steel deck plate Bottom of stiffeners Mid-span of cross-beams

Support ofstiffener

Mid-span ofstiffener

Support Mid-span Top flangeBottomflange

Stress(MPa)

+16.96 –22.38 –55.33 +76.07 –29.43 +36.30

Note + tensile stress

Table 2. Stresses in steel orthotropic deck system


58

Reports


nel suspension bridge (Fig. 11). Theheight of the pylon above the pilecapis 226.14 m, and 160.45 m above thedeck.

The concuk pylon shaft has a boxcross-section with a 1.5 m arc cham-ber. The dimensions of the upper py-lon shaft cross-section in the trans-verse and longitudinal directions are5.5 and 8.5 m respectively, with thick-nesses of 1.25 and 1.0 m. The lowerpylon shaft dimension increases inthe longitudinal direction from 8.5 to11.5 m linearly, and the dimension in

Fig. 10. Local FEM model of anchoragedevice Fig. 11. Configuration of pylon (unit: cm)

Fig. 7. Analysis model of steel orthotropic deck (unit: m)

Fig. 8. Stay-cable anchorage at steel box girder

Fig. 9. Anchorage device



Reports

the transverse direction from 5.5 to9.0 m. Fig. 12 shows the cross-sectionof the upper pylon shaft.

There are two cross-beams be-tween the pylon shafts which are inprestressed concrete with a box sec-tion and 10 m high. The parts of thecross-beam cantilevering beyond thepylon shaft are only for landscapingand decorative purposes.

The foundation beneath the py-lon is in the form of 32 cast-on-sitebored piles with a diameter of 250 cm,which are supported on hard rock.

Piers 38, 40, 41 and 42 have asolid rectangular cross-section withdimensions of 5 m in the transverseand 3.5 m in the longitudinal direc-tion. The foundation is supported oncast-on-site bored piles.

4.5 Stay-cables

The stay-cables are prefabricated par-allel wire cables with a diameter of7 mm and an ultimate tensile strengthof 1770 MPa. The numbers of steelwires are 121, 139, 163, 199, 223 and253. They were tensioned by jacks onthe pylon. The maximum length of thestay-cables is 391.2 m and the maxi-mum weight is 28.4 t. Small dowelsare fitted to the PE of the cable toprevent wind/rain-induced vibrations.

At the time of installation, thestay-cables were jacked to about30 ∼ 35 % of their ultimate capacity.The safety factor of the stay-cables atthe ultimate limit state is about 2.5,which is about 15 % greater than thefactor adopted in Eurocode 1993-1-11.

All cables were considered forthe cable replacement load case. Theobjective of cable replacement is tosatisfy the client that any such futurescenario could be undertaken undersustainable conditions involving lim-ited traffic disruption and without theneed for full closures. The total weightof stay-cables is 1348 t.

5 Construction of main bridge

Since the pylon is located in shallowwater, its foundation was constructedin a steel sheet pile cofferdam (Fig. 13).The pylon was constructed with hy-draulic self-climbing formwork; thecompleted pylon is shown in Fig. 14.

The deck was constructed usingthe double cantilever method; tradi-tional steel travellers were employed

for the steel box girder installation. Foreach segment erection and site welding,the quickest cycle achieved was 4 days.

The deck construction was mon-itored continuously by the on-sitetechnical team to ensure that the re-quired final design vertical and hori-zontal profiles were achieve. The seg-ment erection level was continuouslyadjusted to take into account the dif-ference between theoretical analysisdeflection and site measurements.

The construction of the bridgebegan on 14 February 2005. Deck clo-sure stitches were carried out in therequired sequence by the traveller. Thedeck was finished on 28 March 2008and opened to traffic on 16 Decem-ber 2008. The total deck constructionduration was 47 months. Fig. 15 showsthe completed bridge.

6 Conclusion

The design and construction of thePearl River Northern Channel Bridge,Guangzhou, were completed on timeand to budget as a result of the excel-lent cooperation between client, de-signer, contractor, sub-contractors andsite supervision consultants.

The progress to a successful com-pletion of the project has benefitedfrom having a viable design conceptand solutions that changed very littleduring construction. The considerableinvestment made by the contractor inadditional soil investigations and sci-entific research and tests enabled thedevelopment of a safe and efficient

Fig. 12. Cross-section of upper pylonshaft (unit: cm)

Fig. 13. Steel sheet pile cofferdam

Fig. 14. Completed pylon


60

Reports


substructure and superstructure de-sign.

The bridge has been well receivedby the Guangdong Province govern-ment and the Ministry of Communi-cation of China, who regard the con-struction of the first major cable-stayedbridge in southern China as a symbolof progress and prosperity. It is des-tined to become a significant new land-mark in this important city.

Acknowledgements

The authors would like to acknowl-edge the cooperation and assistance

of the client, vice-general director Mr.Shaojin Zhang, the client’s site repre-sentatives, Mrs. Taike Zhang and Mrs.Hong Cheng, and all the colleagues ofthe design team, the independentchecking team and the constructionteam.

References

[1] BS 5400-3:2000, Steel, concrete andcomposite bridges Part 3: Code ofpractice for the design of steel bridges.British Standard Institution, London.

[2] JTG D62-2004: Code for design ofhighway bridges from reinforced con-

crete and prestressed concretes bridgesand culverts [S], China, 2004.

[3] BS EN 1993-1-10: Design of steelstructures – Part 1-10: Material tough-ness and through thickness properties[S]. CEN, Brussels, 2006.

[4] BS EN 1993-1-5: Design of steel struc-tures – Part 1-5: Plated structural ele-ments, CEN, Brussels, 2005.

[5] BS EN 1993-1-9: Design of steel struc-tures – Part 1-9: Fatigue, CEN, Brussels,2005.

[6] Setra, SNCF: Ponts métalliques etmixtes: Résistance à la fatigue. Guide deconception et de justifications, 1996.

[7] BS EN 1993-1-11: Design of steelstructures – Part 1-11: Design of struc-tures with tension components, 2006.

[8] DIN-Fachbericht 103, Stahlbrücken,Beuth Verlag, March 2003.

Authors:Dr.-Ing. Yingliang Wang, Deputy Chief Engineer of China Railway Major Bridge Reconnaissance & Design Institute Co. Ltd., Wuhan 430056, ChinaDr.-Ing. Chengzao Huang, Department of Communication, Guangdong Province, ChinaIng. Yuncheng Feng, CCCC First Highway Consultants Co. Ltd., Xi’an, China

Fig. 15. The completed main bridge

As one hundred percent, equal sistersLEICHT FRANCE SAS and LEICHTStructural engineering and specialistconsulting GmbH share both the sameshareholder structure and range of ser-vices. General Manager of LEICHTFRANCE is Graduate Engineer MarcelEnzweiler, President is Dr. Ing. LutzSchoene. Chief engineers and contactpersons in Nantes are the membrane experts Sylvain Bernard and JochenArndt.

For more information please visitwww.leichtonline.com.

Leicht France opens in Nantes

Four years after the founding of thecompany, LEICHT is pleased to anno-unce the birth of a French sister com-pany: LEICHT FRANCE SAS saw thelight of day in Nantes and is assumingall the commitments of its German sisteron the French market.

With the establishment of LEICHTFRANCE as a French enterprise, LEICHTintends to bring the successful coopera-tion with French customers and partnersto a new level: the legal basis provided bythis type of stock corporation (Société

News

par actions simplifiée /SAS) not onlysimplifies many business procedures but,in particular, matters related to insurancecoverage. Besides barrier-free communi-cation, the introduction of a French en-gineering team also promises the enrich-ment of LEICHT’s know-how with coun-try-specific expertise.

According to the company, the goal ofLEICHT FRANCE is to establish itselfin the medium term among France’s topexperts with respect to consultation andsupport structure design for projects in-volving ETFE film, lightweight construc-tion and façades.



Reports

reconstructed NSC “Olympic” sta-dium in order to turn it into a first-class location in accordance withUEFA requirements.

2 Conceptual design

The stadium sports a modern archi-tecture: it looks quite organic withinits natural and urban surroundings,and also has a structurally rationaland technologically advanced layoutthat ensures high levels of comfortand safety for its visitors. The stadiumcan accommodate 69 000 spectators,has an area of about 60 000 m2 andits football pitch measures 104 mlong × 72 m wide. In addition, the to-tal built-up area of the stadium coversabout 160 000 m2, the VIP area occu-pies 45 000 m2 and the land improve-ment area extends about 1500 maround the stadium itself (Fig. 2).

1 Introduction

The reconstruction of the “Olympic”National Sports Centre (NSC “Olym -pic”) in Kiev is now underway withinthe framework of a programme beingundertaken in Ukraine to prepare forthe finals of the 2012 European Foot-ball Championship. The UEFA de-cided to include the stadium at theNSC among the four locations in theUkraine where the group, quarter-fi-nal and final matches of the champi-onship are to be played. Unlike someother stadiums, NSC “Olympic” willstill remain a venue for other sportstoo; facilities for track and field eventsare to be created there in accordancewith the requirements of the Interna-tional Association of Athletics Feder-ations. This will make it possible tohost both football matches and inter-national athletics competitions onthe highest level.

The reconstruction of the sta-dium included almost complete dis-mantling of its structures and flood-lighting towers, except for a second tierof stands, in order to reduce the extentand duration of construction works.The whole façade of the stadium is cov-ered by a glass shell, and the standsare covered by a transparent, long-span, suspended membrane/fabricsystem. The 13-storey “Olympic Star”building will be built next to the sta-dium (Fig. 1).

Considerable excavation andconstruction work has to be carriedout around the stadium, including sta-bilizing adjacent slopes to the southand west and constructing an encir-cling concrete gallery to form an ap-proach for spectators. A vast “under-ground city” is to be built beneath thestadium and around it, includingrooms for various purposes and anunderground car park.

The following is a description ofthe basic structural and engineeringconcepts adopted in the design of the

Awaiting the 2012 European Football Championship: some features of the reconstruction of the stadium of the “Olympic” National SportsCentre in Kiev

Aleksandr V. Shimanovsky

DOI: 10.1002/stco.201200008

a)

b)

Fig. 1. A general view of the NSC “Olympic” stadium from a) north and b) west

The article presents results of reconstruction and new construction works at the NationalSports Center ”Olympic“ stadium in the city of Kyiv, which are being performed in Ukraineas preparations for the final tournament of the UEFA European Football Championship2012. It describes conceptual designs and layouts involved. Totaling data are presentedconcerning the number of spectators, the built-up areas, the infrastructure. Structuraldesigns of steel and ferroconcrete constructions are described, results of their compu-tational analysis are included The article also presents data regarding the manufacturingand erection of structures.

13_061-065_SC_shimanovsky (6)_000-000_SC_Bericht (1).qxd 09.02.12 12:39 Seite 61

62

Reports


The general contract for the recon-struction of the stadium was concludedwith JSC Kievgorstroy (Ukraine), andthe contract for the construction ofthe steel roof structure over the standswas awarded to LLC Master-ProfiUkraine Plant. The cables and theiranchorages for the suspended roofstructure over the stadium stands werefabricated by Bridon International Ltd.(UK), the fabric membrane system byHightex GmbH (Germany).

The design of the reconstructionof the NSC “Olympic” was carried outby GMP von Gerkan, Marg und Part-ner (Germany), taking into accountthe UEFA standards that apply tostructures of this kind. The V. Shi-manovsky Ukrainian Institute of SteelConstruction was asked to performan expert appraisal of the design deci-sions for the steel roof structure overthe stands, to devise an erection pro-cedure, to develop a design and tofabricate racks for the assembly ofcolumns and various erection rigging.

In addition to the reconstructionof its bowl, the reconstruction of thestadium included rebuilding the cen-tral grandstand to create a four-levelVIP building with presidential, VIP andcorporate boxes, two-tier sky boxes,bars, restaurants, a mix area, a pressbox, rooms for television, photogra-phers, commentators, security, two sta-tionary and one small intermediate tieron the stand itself, rooms for playersand officials, plus many other plant/ancillary rooms. The “Olympic Star”building will house various functionalrooms for official, sporting and media(a press centre) purposes; there will bea helicopter landing pad on its roof(Fig. 2).

The infrastructure around thestadium is to be fully renovated. Park-

ing will be created for football teams,officials, VIP visitors, media represen-tatives and spectators; separate andsecure passages for evacuation will bearranged. A broad commercial, trad-ing and entertainment area will belaid out in the stadium’s surround-ings. An “Olym pic” electric substation(110 MW) will be constructed near -by.

The old lighting system (based onfour separate free-standing towers)has been dismantled and a new sys-tem installed beneath the roof. It pro-vides the necessary illumination forthe playing field. Two LED indicatorboards have been set up in the sta-dium.

The element involving the great-est responsibility and complexity onthe NSC “Olympic” stadium is theroof structure over its stands, with anoutreach of 65–69 m and an area of48 500 m2. It is one of the world’sbiggest roofs over a sports arena, andat the same time it is a fine combina-tion of architectural aesthetics and op-erational reliability. The features of theroof structure include the following:– The steel columns supporting the

roof are placed outside the stadiumbowl, are not carried on its standsand thus do not transfer any addi-tional loads onto them (Fig. 3a).

– The roof is an oval on plan; it is along-span, suspended, two-layer ra-dial annular structure consisting ofupper and lower radial cables of55–85 mm dia., suspenders and aninner ring (Fig. 3b).

– The roof covering is a fabric mem-brane made of a synthetic materialwith glass-fibre reinforcement anda PTFE coating on both sides; ithas a high strength, high fire resis-tance and high light transparency(Fig. 3c).

– The surface of the roof may experi-ence unavoidable changes in itsgeometry depending on the environ-mental conditions; the process ofshape adaptation is controlled by a

Fig. 2. A general view of the NSC “Olympic” stadium

Fig. 3. The roof structure to the NSC “Olympic” stadium: a) loadbearing frames,b) cable system, c) fabric membrane roof; 1) upper radial cables, 2) lower radialcables, 3) suspenders, 4) internal ring

a)

b)

c)



Reports

computer system which ensuresgood ventilation and prolongednatural illumination of the playingfield, maintains proper temperatureand humidity levels, and eliminatesthe possibility of a “hothouse ef-fect”.

There are about 3000 cables (13–115 mm dia.) in the suspended roofstructure – a total length of 40 kmwhich weighs 765 t. More than 1000clamps are required to connect theseparate cables together. Uniting thecables into a single roof structure re-quires about 22 000 high-precisioncast nodes (Fig. 4).

3 Structural design

The long-span, suspended, two-layer ra-dial annular roof system over the sta-dium’s grandstands hangs on the up-per outer loadbearing ring supportedby 80 steel columns, each 49 m longand weighing about 50 t (Fig. 5).

The upper radial cables supportthe inner ring of the suspended roofsystem, which is made up of a bundleof cables 90 mm in diameter, and hangson the upper outer ring on columnsat a height of about 40 m. The lowerradial cables, attached to the inner

ring and located directly above thestands, hang on the lower outer load-bearing ring supported on columns ata height of about 22 m. The outerloadbearing rings are made of rolledsteel box sections (800 × 1200 mm,wall thickness 30–70 mm).

According to the structural de-sign of the roof, the columns are spacedabout 10.5 m apart and supported viahinges on the reinforced concrete struc-tures of the spectator gallery that en-circles the stadium. The upper load-bearing ring hangs on the tops of thecolumns, attached to those via hinges,in the plane of the cable trusses, andthe lower loadbearing ring is attachedrigidly to the columns. The columns,together with the lower loadbearingring clamped to them, make up a rigidspatial frame that ensures the stabilityof the whole structure, being unitedinto a whole by the prestressed sus-pended roofing system.

The outer loadbearing rings (up-per and lower) undergo compressionbecause of the thrust in the cablestransferred to them; the magnitude ofthe compression can reach 5000 t.The inner ring is in tension – about5500 t. The maximum tensile forcethat the ring is capable of resisting is13 500 t. The steel columns that sup-port the suspended roof have beendesigned for the height of the polygo-nal configuration; the knuckle of itsprofile is at the level of the lower outerloadbearing ring (Fig. 5). The heightof the lower part of the columns isabout 22 m, that of the upper part18 m, which made it possible to reducethe ellipse-like form of the upper outerloadbearing ring in order to decreasethe bending moments that occur thereunder external loads.

The fabric membrane system ispositioned above the lower radial ca-bles of the roof. The system is designedas an “awning” structure with knee-braced props to raise the level of thesurface: this helps to create a pre-stress that stabilizes the shape and al-lows rainwater to drain in a naturalway. The centres of the sections wherethe props are installed will have domeskylights with a transparent coveringof polycarbonate sheets.

The principal structural membersof the roof framework are made ofgrade C355 steel. Their factory con-nections are welded, their on-site con-nections use high-strength bolts.

4 Fabrication and assembly of steelstructures

The contract for the fabrication andassembly of the steel roof structurewas awarded to LLC Master-ProfiUkraine Plant. The installation of theprincipal loadbearing members of theroof (loadbearing columns and outerupper/lower loadbearing rings) hasbeen carried out by JSC Krivorozh-stalkonstruktsia (Ukraine). The in-stallation of the long-span suspendedfabric membrane system was carriedout by Fagioli S.p.A. (Italy), appointedto ship, set up and check hoistingequipment to erect the roof structureto its design position, and by HightexGmbH (Germany), appointed to in-stall the actual fabric membrane.

One of the main project require-ments for the roofing over the standswas to ensure very high precision inthe fabrication and erection of themain loadbearing members. The wholeloadbearing structure for the roof – atotal length of nearly 1 km – had tobe assembled consecutively – elementby element – with minimum tolerances(in some cases set to values lowerthan those in the national standards)in order to guarantee that the closurepoint coincided with the startingpoint of the assembly.

Therefore, the fabricator of theloadbearing structure for the roof, LLCMaster-Profi Ukraine Plant, providedthe production facility with additionalhi-tech metal-processing equipmentand developed specialized technolo-gies for fabricating and supervising themanufacturing and assembly processesto make sure all project requirementswere met. In addition, special rigging

Fig. 4. A cast node for joining the upperand lower radial cables to the inner ring

Fig. 5. Section through and structural design of the suspended roof over the stadiumstands


64

Reports


and assembly benches were set up toachieve the proper quality of the con-struction, assembly and erection. Forexample, the production facility per-formed a check assembly of all largestructural members to form larger sec-tions, then eliminated the discrepan-cies thus found. This satisfied all theproject requirements regarding preci-sion in fabrication and erection of thestructure. Therefore, the project im-plementation procedure included lit-tle or no correction of the loadbear-ing members of the roof after fabrica-tion.

The design of the roof paid spe-cial attention to the technological con-venience of fabrication and the abilityto assemble the loadbearing steel struc-ture to form larger units in order tomake construction simpler and faster.All steel columns were pre-assembledon the ground to form larger units. Theerection of the loadbearing structureconsisted of the following phases: fivecolumns; five sections of the lowerouter loadbearing ring; a section ofthe upper outer loadbearing ring ofthe first span; and so on in this order,taking into account the retardation ofthe erection of sections of the upperouter loadbearing ring by four spans,particularly: first and second columnswith a section of the lower outerloadbearing ring between them; thirdcolumn and a section of the lowerouter loadbearing ring of the secondspan; fourth column and a section ofthe lower outer loadbearing ring ofthe third span; fifth column and a sec-tion of the lower outer loadbearingring of the fourth span, together witha section of the upper outer loadbear-

ing ring of the first span, and so on inthis order until the whole structurewas finished.

Special erection rigging was de-signed for the purpose of installationand for regulating the geometry of theloadbearing structure of the roof. Theworks to be performed on site in-cluded assembling columns from twoshipped items into a single structuralmember; four benches were designedand built for this task (Fig. 6).

The assembled columns wereplaced in the design position one by

one using special travelling towersabout 10.5 m high which, firstly, func-tioned as temporary/erection supportsfor the columns and sections of thelower outer loadbearing ring and, sec-ondly, permitted regulation of the col-umn geometry using special lifting de-vices installed on them (Fig. 7). Afterthe geometry of another group of fivecolumns had been regulated, the trav-elling towers were moved into a newposition to support the next group.Props were added under the columnsalready installed and regulated untilcompletion of the whole loadbearingstructure of the roof (Fig. 8).

After the installation was overand the loadbearing support structurefor the roofing was in place, and itsshape stabilized by a prestress, the ca-ble net of the suspended part of theroofing were erected. For this purpose,each column had a lifting/tensioningdevice attached to it; the devices werecontrolled by a special computer sys-tem.

The installation of the membersof the roof’s loadbearing structurewas performed on the side of the cen-tral (inner) area of the stadium usingvarious lifting plant, including aunique crane, a Demag CC 2800-1,with a capacity of 600 t and a 114 mjib (Fig. 9).

5 Structural analysis

The complexity of the design of theroof over the NSC “Olympic” stadiumand strict safety requirements made itnecessary to perform a sophisticatedanalysis of the behaviour of the roofstructure under most unfavourable

Fig. 6. A general view of the benches for assembling the column segments into larger sections on site Fig. 8. Temporary props under the columns during erection

Fig. 7. A travelling tower for installingcolumns at their design positions andregulating their geometry



Reports

combinations of actions. A designmodel was adopted that included theprincipal loadbearing structures of theroof separated into structural parts.The first of the parts was a loadbear-ing roof framework that consisted ofthe steel columns and the upper andlower outer loadbearing rings (seeFig. 3a). The second was its cable sys-tem, which consisted of the inner ring,the upper and lower radial cables, andsuspenders between those (see Fig. 3b).

The finite element model of theroofing was built by representingnearly all of the loadbearing struc-tures as finite bar elements that resistthe tensile/compressive, bending andtorsion forces. Certain parts of thecolumns and the loadbearing rings(as a rule, in a vicinity of joints) wererepresented as finite plate elements.The model took into account hingedsupports to the loadbearing columns;the hinges were assumed to have twodegrees of freedom in order to permitrotation in two vertical planes. Thetotal number of finite elements in the

model was 5468, with 4880 nodes and20 928 unknowns.

The particular feature of the be-haviour of the roof is that it can resistexternal loads only when the cable sys-tem is prestressed. Therefore, the wholeanalysis consisted of two phases. Phaseone included determining a designconfiguration for the prestressed roof-ing over the stadium using its givenbasic geometry parameters. The initialstate of the roofing was assumed tobe unstressed (both load and prestressset to zero), and the design state ob-tained by a step-by-step application ofthe load and an appropriate reduc-tion in the lengths of the upper andlower radial cables in proportion to acertain parameter α so that the designvalues were reached at α = 1. Phasetwo included an analysis of the roofin accordance with the load cases es-tablished: the dead load of the roofstructure, an operational load causedby auxiliary equipment (lights, electriccables, video screens and their instal-lation rigging, acoustic actions), a pre-

stress load, a wind pressure, a snow/iceload, temperature and seismic actions.Four analyses of the roof were per-formed: distributions of the stressesand strains caused by the prestress andby static loads with a symmetric/asym-metric application of the snow loadwere found; the general stability ofthe structure under the static loadswas verified; natural frequencies andmodes of oscillation were determinedfor the prestressed structure of theroof with and without taking into ac-count the apparent masses of the aux-iliary equipment and snow.

The results of the analysis per-formed for the roof over the stadiumshow that the structure does possessa high loadbearing ability and generalstability. It also has sufficient reservesof strength because the maximumstress in the finite bar elements, evenunder a most unfavourable combina-tion of loads, is 225 MPa (yield point =355 MPa), the stresses in the cablesremain positive and their values arenearly two times lower than the per-missible limits. Secondly, the struc-ture has proved to be hardly sensitiveto uniformly distributed temperatureactions which have little noticeableeffect on the distribution of stressesand strains in the structural membersof the roof. Thirdly, the system pos-sesses a sufficient dynamic stiffnessbecause the frequencies of three low-est natural oscillation modes of theroofing – with prestress, dead load ofstructural members and snow – havethe relatively high values of 0.25, 0.27and 0.29 Hz and are hardly dependenton an actual distribution of snow overthe roof.

Keywords: sports facilities; stadiums;space strcutures; numerical 3D modell -ing; design; manufacturing; assembly

Author:Prof. Dr. Sc. (Eng.) Aleksandr V. Shimanovsky,General director of V. Shimanovsky UkrainianInstitute of Steel Construction, Honoured Workerof Science & Engineering of Ukraine, Academi-cian of Academy of Construction of Ukraine2/1, V. Shimanovsky Street, Kiev, 02660, UkraineE-mail: [email protected]

Fig. 9. Installing a column with the Demag CC 2800-1 and SKG-100 cranes



ECCS news

Conferences

Steel Bridges 2012

The Steel Bridges 2012 – “Sustainablybridging the future: the steel way Designand best practices” will be held in Paris,France on 18–20 April and is organizedby SCMF – Syndicat de la ConstructionMetallique de France. The Conferencewill also host the European SteelbridgesDesign Awards ceremony.

There will be 10 sessions during theconference in which the following topicswill be discussed:

1st Session – The architects and ownerspoint of view

2nd Session – Global cost – Life Cycle3rd Session – Design for sustainability4th Session – Design and calculation5th Session – Steel: the future raw ma-

terial6th Session – Steel bridges – some ex-

amples7th Session – Monitoring and sustain-

ability

Bridges Design Awards

8th Session – Maintenance and repairs9th Session – High Speed Railways

25 years experience10th Session – Bridges of the future

Do not miss this international event inthe heart of the capital Paris and withthe best European specialists in steelbridges.

VIII CMM – Portuguese Confe-rence on Steelwork

On the 24th and 25th of November2011, the 8th edition of the Conferenceon Steel and Composite Constructiontook place in Guimarães, Portugal in“Centro Cultural de Vila Flor”. This eventwas organized by CMM, the PortugueseSteelwork Association and had over 200participants.

Several sessions were held during theconference: “Living with Steel” (ArchitectNeven Sidor from Grimshaw); “Cold-Formed Steel Structures in Seismic Area:Research and Applications” (ProfessorRaffaele Landolfo from the University ofNaples); “Composite Steel-Glass Fins forthe Lobby Façade of Iberdrola Tower”(Eng. Francesc Arbós Bellapart, fromBellapart, S.A.U.); “Construção Metálica:Novas Fronteiras” (Professor AntónioReis from IST); “Best Use of Eurocode 3for Member Design in Frame-Structures”(Professor Richard Greiner, from GrazUniversity of Technology); “MantosMetálicos” (Architect Francisco Vieirade Campos).

Facing the present need of compa-nies to be introduced to CE Marking for Steel products and components soas to fulfill the EN 1090-1 requirements,we emphasize the conference “CEMarking of Structural Steelwork – Im-plications for Manufacturers and Speci-fiers” presented by Dr. Roger Pope,Consultant for the British Construc-tional Steelwork Association and TataSteel Europe), which took place duringthe Conference on Steel and CompositeConstruction.

Some seminars also took place dur-ing the conference: DiSTEEL (Displace-ment Based Seismic Design of STEELMoment Resisting Frame Structures)and Semi-Comp+ (Valorization actionof plastic member capacity of semi-com-pact steel sections – a more economicdesign).

In this 8th edition of the conferencewe also gave the sponsors the opportunityto present their products and services incommercial sessions.

Due to the number of papers submit-ted triple sessions were necessary whichdemonstrates not only the quality of theevent but also the high number of par-ticipants from the both technical andscientific steel community.

The conference also had a technicalexhibition in the foyer of Centro CulturalVila Flor, with the presence of 17 com-panies.

As tradition in the CMM Conferences,a Gala Dinner was held on the first dayof the event. This dinner took place in“Pousada de Santa Marinha”. Lunch wasoffered to the participants in the restau-rant “Café-Concerto”.

Announcements

International events

STEELBRIDGES 2012 – InternationalConference on Steel Bridges18–20 April 2012, Paris, France

NASCC: The Steel Conference 201218–21 April, 2012, Gaylord Texan Convention CenterGrapevine, Texas

ECCS European Steel Day and ECCSAnnual Meeting 201219–22 September, Lisbon, Portugal

**NSCC 2012 – Nordic Steel Construc-tion Conference5–7 September of 2012, Oslo, [email protected]

**ICSA 2013 – 2nd International Con-ference on Structures and Architecture24–26 July 2013, Guimarães, Portugalwww.icsa2013.arquitectura.uminho.pt

6th International Conference on Compos- ites in Construction Engineering (CICE)13–15 June 2012conference.cice2012.it

National events

Connections VII – 7th INTERNA-TIONAL WORKSHOP ON CONNEC-TIONS IN STEEL STRUCTURES30 May to 2 June 2012www.ct.upt.ro/connections/index.htm

1º Congresso Luso – Africano de Construção Metálica SustentávelOrganized by CMM27 July 2012, Luanda, Angola

** 10 % SPECIAL DISCOUNT FOR ECCS INDIVIDUAL MEMBERS

14_066-070_ECCS_News_SC 10.02.12 08:10 Seite 66


ECCS news

Technical Committees (TC) activities

Agenda TC Meetings

TMB – Technical Management BoardChairperson: Prof. L. Simões da SilvaVice-Chairperson: Prof. M. Veljkovic

PMB – Promotional Management BoardChairperson: Mr. Bertrand Lemoine

TC3 – Fire SafetyChairperson: Prof. P. SchaumannSecretary: Prof. Paulo Vila RealDate: 17–18 September 2012, Helsinki,Finland (planned)

TC6 – Fatigue and FractureChairperson: Dr. M. LukicDate: 08–09 March 2012, Berlin, Germany

TC7 – Cold-formed thin-walled SheetSteel in BuildingsChairperson: Prof. J. LangeDate: 28–29 June 2012, (planned)

TWG 7.5 – Practical Improvement ofDesign ProceduresChairperson: Prof. Bettina BruneDate: 28–29 June 2012, (planned)

TWG 7.9 – Sandwich Panels & RelatedSubjectsChairperson: Mr. Paavo HassinenDate: 28–29 June 2012, (planned)

TC8 – Structural StabilityChairperson: Prof. H. H. SnijderSecretary: Dr. Markus KnoblochDate: 15 June 2012, Dresden, Germany

TWG 8.3 – Plate BucklingChairperson: Prof. U. KuhlmannSecretary: Dr. B. BraunDate: 2 March 2012, Stuttgart, GermanyDate: 26 October 2012, Liège, Belgium

TWG 8.4 – Buckling of ShellsChairperson: Prof. J. M. RotterSecretary: Prof. S. Karamanos

TC9 – Execution and Quality ManagementChairperson: Mr. Kjetil MyrheDate: 27 February 2012, Brussels, Belgium

TC10 – Structural ConnectionsChairperson: Prof. F. S. K. BijlaardSecretaries: Mr. Edwin BelderDate: 29 May 2012Date: 4–5 April 2013

TC11 – CompositeChairperson: Prof. R. ZandoniniSecretary: Prof. Graziano LeoniDate: 4 May 2012 Lahti, Finland

TC13 – Seismic DesignChairperson: Prof. R. LandolfoSecretary: Dr. A. StratanDate: 11 May 2012 Budapest, HungaryDate: 23 November 2012, Luxembourg(Planned)

TC14 – Sustainability & Eco-Efficiencyof Steel ConstructionChairperson: Prof. Luís BragançaSecretary: Ms. Heli Koukkari

TC15 – Architectural & Structural DesignChairperson: Prof. P. Cruz

TC News

TWG 7.5 – Practical Improve-ments of Design Guidelines

Chairperson: Prof. Bettina Brune

The last meeting of TWG 7.5 held inIthaca (Cornell University, USA), in cooperation with the European RackingFederation ERF, the American RackManufacturers Institute RMI and guestsfrom Japan. It was an interesting meet-ing dealing with all kind of cold-formedstructures (housing, earthquake design,rack design etc.).

During the meeting there were fivesessions where the following topicswere presented and discussed:

Session 1 – TWG 7.5/REF Meeting (RMI)– Research on Cold-Formed Steel in

Salford, UK– Low and midrise cold-formed steel

building systems in the USA– Theoretical and numerical modelling

of mode-coupling and the yield-arc/yield-eye mechanism

– Benchmark example of rack uprights– Benchmark example of box columns

made of stainless steel– Calibration of an FEM for evaluation

of CFS bolted joints in pitch-roof por-tal frames and MS frames

Session 4 – Joint TWG 7.5/ERF/RMIMeeting– Australian rack specifications and

research – American research on rack systems – Earthquake evaluations from a proba-

bilistic point of view– Seismic design of racks in Japan– ERF-items on Rack Systems– Investigation on the test method for

distortional buckling of compressedpallet rack members

– An Investigation on the design ofsteel storage rack columns using theDSM

– Fatigue of rack systems– Building supervision provisions for

rack structures– Column base tests– Earthquake design according to the

RMI specification

Session 5 – Joint TWG 7.5/ERF Meeting– Seismic resisting cold-formed steel

wall systems– A beginner’s overview of Eurocode –

possible tasks for WG 7.5 to help itsuse easier

– National Annexes to EN 1993-1-3– Evolution Groups for Eurocodes– Lipped channel column in axial com-

pression – Eigenmode imperfectionand strength

The next meeting is expected to be heldon 28–29 June 2012 along with the jointTC7 – meeting.

TWG 7.9 – Sandwich Panels &Related Subjects

Chairperson: Mr. Paavo Hassinen

The European Joint Committee consistsof the Technical Working Group ECCSTWG 7.9 and of Working CommissionCIB W056. The preliminary date of thenext meeting of the Joint Committee is28–29 June 2012. The meeting place isat the moment open.

The Joint Committee has completeda new European Recommendation fordesign of sandwich panels with open-ings. The manuscript is now in linguisticcheck. The Joint Committee has startedto draft two new documents. The firstone will introduce the possibilities tostabilize steel structures such as columnsand purlins against buckling and lateralbuckling by utilizing sandwich panels.The document will be an importantguideline in optimizing the frames andsubstructures of industrial buildings. Thesecond new document will study the ef-fect of claddings which are typicallyfixed in the external face of the wall orroof panels. The document will evaluaterisks and benefits and further, a goodpractice of cladding of sandwich panels.

TWG 8.3 – Plate Buckling

Chairperson: Prof. Ulrike KuhlmannSecretary: Dr. Benjamin Braun

The work of TWG 8.3 concentrates onplate buckling issues in steel platedstructures.

The most recent meeting took placeon 11 November 2011 in Stuttgart, Ger-



ECCS news

many, with presentations and discus-sions on the following subjects:– Proposal of amendments to

EN 1993-1-5– Cross-section classification of angles

(D. Beg)– Modified determination of the critical

column-buckling stress (G. Lener)– Comparison of the effective width

method and the reduced stressmethod (J. Simon-Talero)

– Numerical and experimental analysisof intermediate transverse stiffener instiffened plate girders (F. Sinur)

– Resistance of steel plate girders sub-jected to patch loading (R. Chacon,B. Braun)

– Stress distribution in the flanges ofcorrugated web girders (B. Kövesdi)

– Numerical and experimental analysisof longitudinally stiffened plated gird-ers subjected to bending-shear inter-action (F. Sinur)

Short time outcomeA series of papers on recent topics isprepared for Steel Construction Journal1/2012. They also reflect some plannedamendments for EN 1993-1-5 which willpartly be presented for discussion dur-ing the meeting of CEN/TC250/SC3 inApril 2012.

Next meetingsThe next meetings will be held on 2March 2012 in Stuttgart, Germany, andon 26 October 2012 in Liège, Belgium.

TC 9 – Manufacturing and Erection Standards

Chairperson: Mr. Kjetil Myhre

TC9 is the technical committee withinECCS responsible for making recommen-dations concerning technical specifica-tions for manufacturing and erection. Thecommittee has prepared the publication'Guide to the CE Marking of StructuralSteelwork' to provide practical guidanceon the CE Marking of structural steel-work in accordance with the Construc-tion Products Directive (CPD), the Con-struction Products Regulation (CPR) andthe transition period between CPD andCPR. ECCS acknowledges the adviceprovided by the British ConstructionalSteelwork Association Ltd and access toits publication ‘Guide to the CE Mark-ing of Structural Steelwork’ on whichthis publication is based.

The CE Marking standard for fabri-cated structural steelwork, EN 1090-1:2009 'Execution of steel structures andaluminum structures: Part 1: Require-ments for conformity assessment of struc-tural components' has been cited in theEuropean Commission's Official Journal

(OJ). The Applicability Date is 1st ofJanuary 2011, followed by an 3.5 yearsco-existence period. This means that CEMarking of fabricated steelwork can startand will become mandatory in all Euro-pean countries from 1st of July 2014.

It is hoped that this publication will as-sist the steelwork contractors, their pur-chasing clients and supply chain includ-ing stockholders, part fabricated products(curved steel), proprietary products(purlins, cellular beams etc.) designers,specifiers and construction managers.

The TC9 is also preparing a ‘Check-list for initial inspection and surveillanceof the factory production control FPCaccording to EN 1090-1’. Furthermore,the Committee will develop a websitewith a list of certified fabricators, andprepare a handbook for FPC (FactoryProduction Control).

The next meeting is on the 27th Feb-ruary 2012 in ECCS offices in Brussels. Iwelcome more members to join the com-mittee, please contact your national mem-ber association or send an e-mail topost(at)stalforbund.com

TC 10 – Structural Connections

Chairperson: Prof. F. S. K. BijlaardSecretary: Edwin Belder

The next meeting will be combined withthe next International ConnectionsWorkshop to be held in Timisoara. Themeeting after that will take place on the4 and 5 of April 2013 (location notspecified).

In the 99th meeting held in Karlsruhethe technical committee has been work-ing on the corrigenda and amendmentsfor the EN 1993-1-8.

The following topics were also pre-sented and discussed:Amendments Eurocode 3 part 1-8– Discussion about corrigenda and

amendments on Chapter 7 ofEN 1993-1-8 “Hollow Section Joints

– Approval of the Proposed amend-ments EN 1993-1-8 Comments Ac-tions Decisions

Joints

Presentation of design of joints withfabricated box sections according toEurocode 3.Bjørn Aasen

The introduction of steel grades S420and higher have encouraged Norwegianengineers to use these steels in fabricatedbox sections in various structures. How-ever, practical design rules for structuraljoints seem to be missing, and FEM-analyses very time consuming. Therefore

Bjørn Aasen wants to show the commit-tee that EN 1993-1-8 can be applied forthe design of fabricated box section joints.

New software for the design of hollowsection joints. This is provided for freeby Vallourec & Mannesmann, includ-ing book presenting design tables forstandardized joints.Klaus Weynand, Ralf Oerder

During the last ECCS-meeting in Munichdesign software was presented for com-posite joints. Now a new tool has beendeveloped for the design of hollow sec-tion joints under fatigue loading (seeCidect Design Guide 8

Many tools are available within Eu-rope for open section joint design; how-ever this is not the case for hollow sec-tions. The objective is to provide designtools for standardized joints –?designbook & individual joint layouts – designsoftware.

The first phase will have a limitedscope (most common configurations).This will be approved by German build-ing authorities. The scope shall includethe design rules according EN 1993-1-8,static loading, only RHS sections in ma-terial S355 & S460. The Design Bookshall include design tables and back-ground information for various configu-rations (K, N, X, Y, T) in S355 and S460.

A copy of the design book will beprovided to the members at next ECCSTC-10 meeting.

NA to EN 1993-1-8, some commentsabout National Annexes regardingpractitioners.

Work of CEN TC 185WG6 in particu-lar the difficulties encountered to up-date the EN 14399-1 standard for pre-loaded “Structural Bolting” on theevolution of the EN bolt standards,bolts. (Ivor Ryan & Maël Couchaux)The standard series EN 14399; Bolts forPreloading has been expanded up to 10parts since 2005. Part 1 covers the CEmarking (requirements). The EN 14399-1needs modification in order to extendthe scope of CE marking to parts 7–10.However, this encounters problemswith the CEN expert. Ivor Ryan madesome remarks about the performancerequirements of part 2 and the actualbehavior under tightening. Informationfrom a manufacturer indicated that typeHV had difficulties meeting Δθ1 require-ments and sometimes Δθ2 requirements.Ivor Ryan concluded that it is question-able whether a HV product (as describedin part 4) is suitable for tightening bythe combined method unless it meetsthe now “optional” requirement onΔθ1min during the tightening test.



Composite joints

The experimental behavior of heatedcomposite joints, subject to variablebending moments. (Aldina Santiago)Within the European RFCS RobustfireProject research is done to obtain newdesign criteria of car parks with suffi-cient robustness under localized fire.

The tested sub-structure (experimen-tal test) was a part of an open car parkbuilding. Seven experimental tests weredone with various high temperatures.Ceramic pad heating elements wereused to reach these temperatures. Bend-ing moments were applied and testedwith/without axial restraint. Failuremode was concrete crushing and defor-mation of the steel end-plate at centreand bottom part. Other failure modeswere local buckling of the bottom flangeand fractured bolts (M30 bolts-10.9).

Connectors

Tests on connections with injectionbolts with high bearing stresses (NolGresnigt)In Amsterdam, many M27 injectionbolts are applied in a the glass roof ofan extension of the Central Railway Sta-tion. Because of the glass, rather tightlimits were put on the deformations ofthe supporting steel structure. Highstrength friction grip bolts were not agood solution because of the necessityto take away the paint coating of thesteel at the contact surfaces and thedanger of corrosion pollution later on(the steel remains visible for the public).

It was asked whether for the load casewith wind in the ultimate limit state,higher bearing stresses could be allowedthan the recommended values that arebased on long duration creep tests. Thebearing stresses in this ULS load casecould be more than twice the recom-mended values.

Recent test results by Prof Darko Begindicated that indeed for short durationloads higher bearing stresses can be al-lowed. In Delft additional short durationcyclic tests were carried out, some withmore than double the recommendedbearing stresses. The test results togetherwith test results from research in Ljubl-jana were sufficient to allow the muchhigher bearing stresses for 90 % of theglass roof in Amsterdam. Only at theedges of the roof where wind gusts aremost severe, additional strengtheningwas necessary.

A. Load indicated washers.Frans Bijlaard informs the committeeabout the latest developments on DTIwashers. According EN 1090-2 no morethan 10 % of the indicators are allowed

ECCS news

to be fully impressed. However, the fab-ricators want to change this rule (checkon over-tightening) by making a newrule in EN 14399, and declare it bind-ing.

This could very well endanger thesafety of the connection accordingEN 1090-2.

Bjørn Aasen, also member of TC135,will make attempts to discuss about thisissue within TC135 once again.

Bolts

Document Tragverhalten von Scher-Lochleibungsverbindungen mit demBeginn des Gewindes in der Nähe derScherfuge.Part of document:In bolted connections the shear load iscalculated with the cross section of thebolt. It is important whether the shankor the thread is in the shear plane. Inthe new code for bolts DIN EN 14399-4the dimension of the bolts are modified.Compared to the previous code, theDIN 6914, in many situations the threadis closer to the shear plane or even inthe shear plane. In the current standardsthere is no requirement about how longthe shank has to protrude into the shearplane.

The question is if such a rule shouldbe added to the code. Should engineerscheck bolt lengths on drawing or is itthe responsibility of the steel construc-tion fabricator?

Klaus Weynand feels that this is theresponsibility of the fabricator to act ac-cording the design of the joints.

High strength steel

A. Current research focuses at KITThomas Ummenhofer presented variousresearch projects which are currently inprogress at KIT.

These projects include the optimiza-tion of supporting structures for offshorewind energy convertors using HIFITand the development of a joining tech-nique for adhesive bonding of hollowsections.

B. Effects of post treatment on the fa-tigue life of ultra high strength steelsPhilippWeidner presented research con-cerning Domex 960 steel treated withHIFIT. Especially in crane design weightsmust be as low as possible. Thereforehigher strength steels are used. The testresults show that the fatigue life in theLCF range can be extended 5–6 timesusing HIFIT-treatment. All specimenfailed at the weld toe. Philipp Weidnerconcluded that ultra high strength steelsoffer more potential with regard to fa-tigue compared to lower yield steels.

Publications

ECCS book “Design of joints”, to beconsidered a background-document toEN 1993-1-8 (new book under develop-ment) (Klaus Weynand, Jean-Pierre Jas-part, Jürgen Kuck)

Request for information from KlausWeynand, regarding books for the de-sign of connections according EC. Thiscould also be documents or books withdesign tables. Books like BCSA-simplejoints (green book). Work on the bookis well underway. Inform KWor JPJabout status of their books and con-tents, so they can refer to these booksand documents also. Please reply tothem before end of May.

These meetings were a good opportu-nity for delegates to present their re-search, recent projects and make sugges-tions to improve the EN-code based onresearch or practice.

TC 11 – Composite Structures

Chairperson: Prof. Riccardo ZandoniniSecretary: Prof. Graziano Leoni

Before the last meeting held in Timisoaraon November 4th, 2011, Prof. Jiri Stud -nicka has resigned as Full Member inthe TC 11. He will continue his involve-ment in the TC activities as a Corre-spondent Member. In this way the TCwill not loose his important input as hecontinues his active participation, simi-larly to other Emeritus Professors suchas Jean Marie Aribert, Roger Johnson,and Joël Raoul, who enhance the activi-ties of the TC with their experience andexpertise.

Two new members joined the TC 11:Jakub Dolejs from Czech Republic, whoreplaces Prof. Jiri Studnicka, and ManuelRomero from Spain. Nineteen Full Mem-bers and seven Corresponding Members,from thirteen European countries, Aus-tralia and New Zeeland, currently con-stitute the TC 11.

Besides the ongoing activities aimedat preparing documents that shouldcomplement Eurocode 4 with referencesto issues on shear connections, flooringsystems and composite frames, othertopics are currently being discussedwithin TC 11.

In the last meeting, Prof. Jean-PaulLebet, from the Ecole PolytechniqueFédérale de Lausanne, and Prof. DennisLam, from the University of Bradford,reported about advances in research oncomposite bridges and elliptical hollowcolumns, respectively.

Due to its valuable members, TC 11represent a permanent and useful sourceof design information for compositeconstruction in Europe.


The next meeting of ECCS-TC 11 willbe held in Lahti (Finland) on 4th May2012.

TC 13 – Seismic Design

Chairperson: Prof. R. LandolfoSecretary: Prof. A. Stratan

In 2011 TC 13 organized two meetings:the former in Rennes (France) on 20th

May, the second in Ljubljana (Slovenia)18 November. The main activities of TC13 concentrated on the seismic designof steel structures with special regard tothe following aspects:

1. TC 13 commentary and integrationto EN 1998-1: 2004: on the basis of awork document developed in the lastyear within TC 13 activity, it is underpreparation the following ECCS pub-lication: “Assessment of EC 8 provi-sions for seismic design of steel struc-tures”. In this book the main issuesin EN 1998-1: 2004 that need clarifi-cation and/or further developmentare described and discussed. Thedocument is organized in twelve Sec-tions and one Annex.

2. TC 13 dissemination activity: Withinthe “STESSA 2012” conference (“Be-haviour of Steel Structures in Seis-mic Areas”), held in Santiago (Chile)on the last January 9–11 and alsosupported by ECCS, a special sessiondedicated to TC 13 activity was orga-nized. In that occasion 8 papers con-cerning the main topics dealt witheach TC13 TWGs were presented.

3. Next meeting: The next TC 13 meet-ing is fixed for 11 May in 2012 in Bu-dapest. The following one is plannedfor 23 November 2012 in Luxembourg.

New Publications

Guide on the CE Marking ofStructural Steelwork

The CE Marking standard for fabricatedstructural steelwork, EN 1090-1: 2009


‘Execution of steel structures and alu-minum structures: Part 1: Requirementsfor conformity assessment of structuralcomponents’ has been cited in the Euro-pean Commission’s Official Journal (OJ).The Applicability Date is 1st of January2011, followed by an 18 months co-exis-tence period. This means that CE Mark-ing of fabricated steelwork can start andwill become mandatory in most Euro-pean countries from 1st of July 2012.CE Marking is optional in the UK, Re-public of Ireland, Finland, Sweden andNorway for one more year, until 1st ofJuly, when the Construction ProductsDirective (CPD) is replaced by the fullConstruction Products Regulations(CPR).

TC 9 is the technical committeewithin ECCS responsible for makingrecommendations concerning technicalspecifications for manufacturing anderection. The committee has preparedthis publication to provide practicalguidance on the CE Marking of struc-tural steelwork in accordance with theConstruction Products Directive (CPD),the Construction Products Regulation(CPR) and the transition period betweenCPD and CPR.

It is hoped that this publication willassist the steelwork contractors, theirpurchasing clients and supply chain in-cluding stockholders, part fabricatedproducts (curved steel), proprietaryproducts (purlins, cellular beams etc.)designers, specifiers and constructionmanagers.

Thin Walled Structures – Recent Research Advancencesand Trends (2 volumes)

The Proceedings of ICTWS 2011 – RECENT RESEARCH ADVANCESAND TRENDS, organized in two vol-umes, contain 9 keynote lectures pre-sented by outstanding scientists and115 papers, by authors from 33 coun-tries of Africa, Asia, Australia, Europe,North America and South Africa.These papers are grouped in 9 maintopics i.e.:

Session 1: Buckling;Session 2: Post-bucking analysis and

failure modes;Session 3: Behaviour of thin-walled

structures under extremeloadings;

Session 4: Connection on thin-walledstructures;

Session 5: Cold-formed steel structures; Session 6: Composite structures;Session 7: Storage racking;Session 8: Shell and space structures;Session 9: Plated Structures.

Full price: 95 €Available at ECCS online Bookstore(www.steelconstruct.com). Limited toexisting stock.

Projects

SKILLS – Steel Construction IndustryLive Long Learning Support

Skills is a new project funded by theLEONARDO DA VINCI Lifelong learn-ing programme in France and is beingled by CTICM (Centre Technique Indus-triel de la Construction Mátallique), incollaboration with PIKS and ASCEM,the steel fabricator associations in Polandand Spain, Warsaw University of Tech-nology, the Universitat Politècnica deCatalunya and ConstruirAcier, the steelconstruction promotional organizationin France.

The aim of SKILLS is to assist trainingproviders in France, Poland and Spainmeet the substantial needs of the steelconstruction industries in these coun-tries, arising from the changes in work-ing practices due to implementation ofthe Eurocodes. The project will deliverpractical guidance which will demon-strate Eurocode application in steelbuilding design.

The tangible outcomes will be train-ing materials, in tried and tested formats,aligned to national priorities and in na-tional languages, for delivery by nationalteachers and trainers.

ECCS news



The international journal “Steel Construction – Design and Research”publishes peer-reviewed papers covering the entire field of steel con-struction research and engineering practice, focusing on the areas ofcomposite construction, bridges, buildings, cable and membrane struc-tures, façades, glass and lightweight constructions, also cranes,masts, towers, hydraulic structures, vessels, tanks and chimneys plusfire pro tection. “Steel Construction – Design and Research” is the en-gineer ing science journal for structural steelwork systems, which em-braces the following areas of activity: new theories and testing, design,analysis and calculations, fabrication and erection, usage and conver-sion, preserving and maintaining the building stock, recycling anddisposal. “Steel Construction – Design and Research” is thereforeaimed not only at academics, but in particular at consulting structuralengineers, and also other engineers active in the relevant industriesand authorities.

“Steel Construction – Design and Research” is published four times ayear.

Except for manuscripts, the publisher Ernst & Sohn purchases exclu-sive publishing rights. Ernst & Sohn accepts for publication only thoseworks whose content has never appeared before in Germany or else-where. The publishing rights for the pictures and drawings madeavailable are to be obtained by the author. The author undertakes notto reprint his or her article without the express permission of the pu-blisher Ernst & Sohn. The “Notes for Authors” regulate the relation -ship between author and editorial staff or publisher, and the compo-sition of articles. “Notes for Authors” can be obtained from the pub -lisher or via the Internet at www.ernst-und-sohn.de/zeitschriften.

The articles published in the journal are protected by copyright. Allrights are reserved, particularly those of translation into foreign lan-guages. No part of this journal may be reproduced in any form what-soever without the written consent of the publisher. Brand-names ortrademarks published in the journal are to be considered as protectedunder the terms of trademark protection legislation, even if they arenot individually identified as such.

Manuscripts are to be sent to the editorial staff.

If required, offprints or run-ons can be made of single articles. Requestsshould be sent to the publisher.

Current pricesThe journal “Steel Construction – Design and Research” comprisesfour issues per year. In addition to “Steel Construction – Design andResearch print”, the PDF version “Steel Construction – Design andResearch online” is available on subscription through the “Wiley On-line Library” online service.

Members of the ECCS – European Convention for ConstructionalSteelwork receive the journal Steel Construction as part of their mem -bership.

Prices exclusive VAT and inclusive postage. Errors and omissions ex-cepted. Subject to change without notice.

Prices are valid from 1st September 2011 until 31st August 2012.

Personal subscriptions may not be sold to libraries nor used as librarycopies.

A subscription lasts for one year. It can be terminated in writing atany time with a period of notice of three months to the end of thesubscription year. Otherwise, the subscription extends for a furtheryear without written notification.

Bank detailsDresdner Bank Weinheim, A/C No: 751118800Bank sort code: 67080050, SWIFT: DRESDEFF670

Periodical postage paid at Jamaica NY 11431. Air freight and mailingin the USA by Publications Expediting Services Inc., 200 MeachamAve., Elmont NY 11003, USA

POSTMASTER:Send changes of address to“Steel Construction – Design and Research”c/o Wiley-VCH, 111 River Street, Hoboken, NJ 07030, USA

Prices print print + online single issue

personal 154 € 178 € 44 €

personal 256 sFr 295 sFr 73,60 sFr

personal 243 $ 279 $ 69,86 $

institutional 520 € 598 € 149 €

institutional 865 sFr 995 sFr 248,69 sFr

institutional 818 $ 941 $ 235,18 $

PublisherWilhelm Ernst & SohnVerlag für Architektur und technische Wissenschaften GmbH & Co. KGRotherstrasse 2110245 BerlinGermanyPhone +49 (0)30 47031-200Fax +49 (0)30 [email protected]

Journal for ECCS membersThe ECCS – European Convention for Constructional Steelwork and Ernst & Sohn have agreed that from 2010 onwards Steel Construction is the official journal for ECCS members. www.steelconstruct.com

Editorial staffEditor-in-chief: Dr.-Ing. Karl-Eugen Kurrer, Ernst & SohnPhone +49 (0)30 47031-248, Fax -270E-mail: [email protected]

Editorial boardChair: Luis Simões da Silva (Portugal)

Advertising managerNorbert Schippel, Ernst & SohnPhone +49 (0)30 47031-252, Fax -230E-mail: [email protected]

Services for customers and readersWiley-VCH Customer Service for Ernst & SohnBoschstrasse 12, D-69469 WeinheimTel.: +49 (0)800 1800 536 (with Germany)Tel.: +44 (0)1865476721 (outside Germany)Fax: +49 (0)6201 [email protected]: www.wileycustomerhelp.com

Layout and typesettingLVD I BlackArt, Berlin

ProductionNEUNPLUS1 GmbH, Berlin

© 2012 Ernst & Sohn Verlag für Architektur und technische Wissenschaften GmbH & Co. KG, Berlin

Inserts in this issue:

Surveyor Events, SW1V 255 London, UK;

Verlag Ernst & Sohn GmbH & Co. KG, 10245 Berlin

• Darko Beg (Slovenia)• Frans Bijlaard

(The Netherlands)• Luís Bragança (Portugal)• Dinar Camotim (Portugal)• S. L. Chan (P. R. China)• Paulo Cruz (Portugal)• Dan Dubina (Romania)• László Dunai (Hungary)• Morgan Dundu

(Rep. of South Africa)• Markus Feldmann (Germany)• Dan Frangopol (USA)• Leroy Gardner (UK)• Richard Greiner (Austria)• Jerome Hajjar (USA)• Markku Heinisuo (Finland)• Jean-Pierre Jaspart (Belgium)• Ulrike Kuhlmann (Germany)• Akimitsu Kurita (Japan)• Raffaele Landolfo (Italy)

• Guo-Qiang Li (P. R. China)• Richard Liew (Singapore)• Mladen Lukic (France)• Enrique Mirambell (Spain)• Kjetil Myhre (Norway)• Kim Rassmussen (Australia)• John Michael Rotter (UK)• Peter Schaumann (Germany)• Gerhard Sedlacek (Germany)• Bert Snijder

(The Netherlands)• Thomas Ummenhofer

(Germany)• Ioannis Vayas (Greece)• Milan Veljkovic (Sweden)• Pedro Vellasco (Brazil)• Paulo Vila Real (Portugal)• Frantisek Wald

(Czech Republic)• Riccardo Zandonini (Italy)• Jerzy Ziółko (Poland)

20_Impressum_Vorschau_1-12_16_SC_Impressum_1-09.qxd 10.02.12 08:20 Seite 1

Preview

Steel Construction 2/2012

(subject to change without notice)

nn Yes, we would like to order Steel Construction – Design and Research.beginning with the next issue for one year at first. The subscription can be stopped three month before the reference period ends. If we do not cancel our subscription, please deliver for one additional year.

Please choose: ↓ or ↓Prices print print + online

personal / company 154 € 178 €

personal / company 256 sFr 295 sFr

personal / company 243 $ 279 $

institutional / universities 520 € 598 €

institutional / universities 865 sFr 995 sFr

institutional / universities 818 $ 941 $

nn We would like to receive more information about individual membership of the ECCS – European Convention for Constructio-nal Steelwork (www.steelconstruct.com).

nn Please send us a free sample copy / 1 issue of the journal Stahlbau (published in German)

Invoice and delivery address:

Company VAT-No.

Titel, Last name, First name Street/No

Department Country, Zipcode, City

email Phone

We guarantee you the right to revoke this order within two weeks, please mail to Verlag Ernst & Sohn, Wiley-VCH, Boschstr. 12, D-69469 Weinheim

All prices are net-prices exclusive of VAT but inclusive postage and handling charges.Date Signature Prices ar valid from 1st September 2011 until 31st August 2012.

Steel ConstructionDesign and Research

– Gerhard Sedlacek, a passionate engineer– EN 1990 – innovative head Eurocode– Plate buckling to EN 1993-1-5– Steel seismic upgrades for RC structures– Innovative steel connections– Capacity analysis of Möller girders– Structural design of TCI building, Bucharest– Kolomoen Bridge, Norway– “Olek” Zienkiewicz – FEM pioneer– Jörg Schlaich – structural artist and teacher

4Volume 2December 2009ISSN 1867-0520

7

Fax +49 (0)30 47031 240

journal for ECCS members

Markku Heinisuo, Henri Perttola,Hilkka RonniComponent method for end plate joints,modeling of 3D frames: literature review

Literature of modelling steel frames in 3Dusing beam finite elements is presented.The development of member modellingis described first followed by joint mo-delling. The review ends with a brief in-troduction of the 3D component modelfor end plate joints. The next part inclu-des three examples of the 3D componentmethod: base bolt joint, beam-to-columnjoint and member splice joint.

Johan Maljaars, Frank van Dooren,Henk KolsteinFatigue assessment for deck plates inorthotropic bridge decks

Various bridges in highways suffer fromfatigue cracks in the deck plate of ortho -tropic steel bridges. This article providesan assessment procedure for the fatiguelifetime of the deck plate. The lifetimedetermined with the assessment pro -cedure agrees well with the lifetime ob-served in practice. The required deck

plate thickness for new bridges is deter-mined with the assessment procedure. Amethod is developed to quantitativelyaccount for results of inspections in thedesign.

Matthias Wieschollek, Nicole Schillo,Markus Feldmann, Gerhard SedlacekLateral torsional buckling checks ofsteel frames using 2nd order analysis

This article summarizes rules from EN1993-1-1 to perform a 2nd order analy-sis for the flexural and lateral torsionalbuckling check of members and systemsthat are loaded in the strong plane oftheir cross-section and shall not fail byout-of-plane instability.The rules relevant in EN 1993-1-1 arerules for equivalent geometrical imper-fections resulting from the elastic buck-ling modes with amplitudes derivedfrom column buckling tests according toEN 1990 — Annex D and rules for the li-near superposition of utilization rates inthe strong plane and the weak plane asused for the evaluation of the tests todetermine the amplitude of imperfec-tions and the partial factors.

To determine the elastic buckling modesin general computer-assistance is requi-red; it gives both modal out of plane dis-placements but also modal out of planebending moments for the relevant flan-ges of the profiles. The use of out ofplane bending moments makes the as-sessment clear and simple. It leads to amagnification factor for the in-plane-uti-lization rate that allows to detect the de-sign point x = xd along the length of themember or system where the cross-sec-tional check is relevant. It also showsthat the results of the 2nd order analysisfor this design point and the use ofbuckling curves for member checks atthis design point are identical.The assessment procedure with 2nd orderanalysis is generic and applicable to alldesign cases whereas the member checkswith buckling curves as specified in EN1993-1-1 are intended to be valid for spe-cific loadings and boundary conditionsonly (object oriented rules). They shouldbe consistent with the generic rules. Inthis respect this article is a contributionfor the further evolution of EN 1993-1-1.

20_Impressum_Vorschau_1-12_16_SC_Impressum_1-09.qxd 09.02.12 12:41 Seite 2

ww

w.e

rnst

-un

d-s

oh

n.d

e/Z

eit

sch

rift

en



Kundenservice: Wiley-VCHBoschstraße 12D-69469 Weinheim


Steel Construction unites in one journal the holistic ap-

proach to steel construction. In the interests of “construction

without depletion”, it skilfully combines steel with other forms

of construction employing concrete, glass, cables and mem-

branes to form integrated steelwork systems. The scientific and

technical papers in Steel Construction are primary publications.

This journal is aimed at all structural engineers, whether active

in research or practice.

All papers will be subjected to a rigorous peer review process

prior to acceptance for publication via ScholarOne Manuscripts.

The ECCS - European Convention for Constructional Steelwork -

and Ernst & Sohn have agreed that the journal Steel Construc-

tion founded by Ernst & Sohn in 2008 is to be the official jour-

nal for ECCS members from 2010 onwards.

For further information and journal guidelines please contact:

Steel Construction

Wilhelm Ernst & Sohn Verlag für Architektur und

technische Wissenschaften GmbH & Co. KG

Rotherstr. 21

10245 Berlin

Germany

[email protected]://mc.manuscriptcentral.com/stco

Volume 5, 2012 / 4 issues per year.

Editor-in-chief: Dr.-Ing. Karl-Eugen Kurrer

For further information, please visit:

subscription:

www.ernst-und-sohn.de/steel-construction

membership:

www.steelconstruct.com

Call for Papers

Fachwissen ohne Grenzen

0163

4100

16_d

p

Journal for ECCS members –European Convention forConstructional Steelwork

21_SC_U3_1-12_SC 09.02.12 12:42 Seite 1

* -P

rices

are

val

id in

Ger

man

y, e

xclu

sive

ly, a

nd s

ubje

ct t

o al

tera

tions

. Pric

es in

cl. V

AT.

Boo

ks e

xcl.

ship

ping

. Jou

rnal

s in

cl. s

hipp

ing.

019

6300

006_

dp

On

lin

e-O

rde

r: w

ww

.ern

st-u

nd

-so

hn

.de





Bauingenieure mögen Fußwegbrücken mäßiger Breite und Traglasten weniger beachten als Straßen- und Eisenbahnbrücken von spektakulärer Spannweite und Konstruktion. Städtebauer und Landschaftsplaner hingegen erachten innerstädtische Rad- und Gehwegbrücken für weitaus entscheidender und veranlassen in der Regel Wettbewerbe. Da es keine einheitlichen Entwurfsrichtlinien gibt, bilden die Erfah-rungen aus realisierten Brücken eine wichtige Informationsquelle für Planer.

Das vorliegende Buch enthält über 100 Beispiele, die in den letzten zehn Jahren weltweit gebaut wurden: offene Fuß- und Radwegbrücken, Viehtrieb- und Medien-brücken sowie einige geschlossene Verbindungsstege.

Die Beispielsammlung ist nach Tragwerkstypen und Spannweiten gegliedert. Zu jeder Brücke gibt es eine kurze Darstellung der Randbedingungen und eine Bauwerksbeschrei-bung, illustriert anhand von Fotos, Grund- und Aufrissen, besonderen Konstruk tionsdetails. So-mit bildet es eine Fundgrube für Bauingenieure.

Structural engineers are often more interested in road and rail bridges of spectacular construction and enormous spans than in relatively narrow footbridges built for modest loads. City planners and landscape architects, on the other hand, see inner-city pedestrian and cycle bridges as important architectural elements and generally invite bridge builders to compete to find the winning design. As there are no official guidelines for the design of footbridges, the building techniques and performance of existing bridges are an important source of information for planners.

This book contains over 100 examples of bridges built worldwide over the last ten years: open pedestrian and cycle bridges, cattle and utility bridges and some enclosed skyways. The collection is arranged according to load-bearing structure. For each bridge there is a brief description of the location and the structural system, including special construction details, and illustrated by photographs, plans and elevations. “Footbridges“ is a treasure trove for structural engineers.

Literature for bridge building by Ernst & Sohn

When bridges fail, often with loss of human life, those involved may be unwilling to speak openly about the cause. Yet it is possible to learn from mis-takes. The lessons gained lead to greater safety and are a source of innovation.

This book contains a systematic, unprecedented overview of more than 500 bridge failures assigned to the time of their occurrence in the bridges‘ life cycle and to the releasing events. Primary causes are identified. Many of the cases investigated are published here for the first time and previous in-terpretations are shown to be incomplete or incorrect. A catalogue of rules

that can help to avoid future mistakes in design analysis, planning and erection is included.

A lifetime‘s work brilliantly compiled and courageously presented – a wealth of knowledge and experience for every structural engineer.

J O A C H I M S C H E E R

Failed BridgesCase Studies, Causes and Consequences

2010. 321 pages. 120 fi g. 15 tab. Hardcover.

79,–*ISBN 978-3-433-02951-0

K L A U S I D E L B E R G E R

The World of FootbridgesFrom the Utilitarian to the Spectacular

2011. 183 pages, 351 fi gures. Softcover.

69,–*ISBN 978-3-433-02943-5

K L A U S I D E L B E R G E R

Fuß- und RadwegbrückenBeispielsammlung

2011. 182 S., 351 Abb., Br.

49,90*ISBN 978-3-433-02937-4

22_SC_U4_1-12_SC 20.02.12 13:52 Seite 1

Documents

Steel Construction 01-2012 (sample issue)