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Structural Engineering International

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Shaft Design for the Rijnlandroute Bored Tunnel

Hans Mortier Civ. Eng., Engineering Manager EPC Projects & Bart PeerdemanCiv. Eng., Associate Director

To cite this article: Hans Mortier Civ. Eng., Engineering Manager EPC Projects & Bart PeerdemanCiv. Eng., Associate Director (2020): Shaft Design for the Rijnlandroute Bored Tunnel, StructuralEngineering International, DOI: 10.1080/10168664.2020.1759389

To link to this article: https://doi.org/10.1080/10168664.2020.1759389

Published online: 03 Aug 2020.

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Shaft Design for the Rijnlandroute Bored TunnelHans Mortier, Civ. Eng., Engineering Manager EPC Projects, Design Department, DEME Infra Marine Contractors (DIMCO),

Zwijndrecht, Belgium; Bart Peerdeman, Civ. Eng., Associate Director, Infrastructure Rotterdam, Royal Haskoning DHV, Rotterdam,

The Netherlands. Contact mortier.hans@deme-group.com

DOI: 10.1080/10168664.2020.1759389

Abstract

This article elucidates how thelaunching and arrival shafts of theRijnlandroute bored tunnel, nearLeiden in the Netherlands, aredesigned and executed. Specialattention is given to the behavior ofthe diaphragm walls during excavationof the shaft and when the tunnelboring machine enters the surroundingsoil mass. The global stability of theshafts, subjected to high asymmetrichorizontal loads, and an executionmethodology ensuring watertightmining through the diaphragm wallsare the main points of interest. Theadoption of a permanent bell as analternative to the more commonlyused starting plug will be described, aswell as the design details of the “softeye” to ensure a good concretingprocess of these diaphragm walls.

Keywords: GFRP reinforcement;launching shaft; arrival shaft; diaphragmwall; global stability

Introduction

The Rijnlandroute is the new linkbetween the existing A4 and A44 high-ways in the delta region near the city ofLeiden in the Netherlands (Fig. 1). Thenecessity for this new link originatedfrom the large number of daily commu-ters seeking a traffic-jam-free passagethrough the heart of the city. Theproject comprises a 2.2 km twin-tube11 m diameter bored tunnel intercon-nected by eight cross-passages, anopen cut of 1.4 km with an aqueduct,and two cut-and-cover entrances ofapproximately 300 m each. On top ofthis, two major interchanges compris-ing two 300 m long concrete box-girder bridges and several otherbridges and underpasses are alsoincluded in the project. More detailedinformation regarding the boredtunnel part can be found in Ref. [1].

Tender documents included a refer-ence design, elaborated by the client,showing a slightly shorter bored

tunnel length than the tunnel in thefinal design. As the launching andarrival shafts were quite deep andtherefore costly, cut-and-cover sectionsdue to the soft upper soil layers neededto be made, the joint venture to whichthe contract was awarded decided tolengthen the mechanized tunneldrives as much as possible. The uppersoil layers, down to a depth of−22.30 m normal Amsterdam level(NAP), consist of clayey and peatysoil types, having specific weights of17 kN/m3, internal friction angles ϕranging from 17.5 to 27.5° and cohe-sion values cu of 7 kPa. The underlyingPleistocene sand layer has a specificweight of 19 kN/m3, an internal frictionangle ϕ of 35° and no cohesion.

The requirement of a maximumspacing between consecutive cross-pas-sages of 250 m limited the totallengthening to 57 m. Horizontal align-ments were adapted as well to ensurea small but feasible spacing betweenthe two main tunnels at the interfaceswith both shafts. The contract awardwas not only based upon the submittedpricing offer, as virtual price reductionswere granted to the bidders if theycomplied with well-defined criteria.One of these criteria was the reductionof hindrance towards local residentsduring the construction period. Oneof the important promises putforward by the contractor to complywith this criterion was the totalabsence of pile driving when installingretaining walls and pile foundations.All foundation piles and temporary

retaining walls were installed usingvibrational energy. This meant that nocombined retaining walls (sheet pileswith open-ended tubes or H-profiles)were possible and the structuralcapacity of the retaining walls waslimited to the implementation of high-strength (steel grade 52) AZ-52 sheetpile profiles. One single strut oranchor level was preferred, or evennecessitated in the case of the shafts,as no intermediate level in front ofthe tunnel boring machine (TBM)could be tolerated. These circum-stances led to the choice of inclinedanchors from section 3 and upwards,a strutting frame in section 2 anddiaphragm walls in section 1 (shaft)(Fig. 2). Adopting inclined anchorsfacilitated the lowering of the back-upgantries of the TBM, but in section 2a stiffer strut frame support wasnecessary to make the AZ-52 ade-quate. The significantly larger depthof the shafts, due to the circular formof the TBM, was one of the reasonsfor adopting diaphragm walls.

Constraints of the ShaftDesign

The length of the launching shaft isdictated by the space needed to lowerthe TBM parts and assemble them infront of the bell. The bell is the struc-ture that enables the excavationchamber of the TBM to be pressurizedbefore the frontal diaphragm wall ismined through. Part of the frontal

Fig. 1: Overall view of the new N434 link

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wall will be mined through by the TBMand therefore needs to comprise glassfiber reinforced polymer (GFRP)reinforcement instead of ordinarymild steel reinforcing bars (rebars),the so-called “soft eye”. Both struc-tures, the bell and the soft eye, as wellas the tympan wall, are indicated inFig. 9. Their respective uses will beexplained in the following paragraphs.Apart from horizontal soil and waterpressures and vertical loads, the dia-phragm walls need to cope with theloads induced by the starting or arriv-ing TBM. As the shallow sections ofthe ramp are made using dewateringinstead of underwater concrete, thereis also an unfavorable unbalance inthe water pressures. Break-ins (fromthe launching shaft into the soilmass) and break-outs (out of the soilmass into the arrival shaft) requiremeasures to ensure water tightnesswhen tunneling under the existingwater level. Currently adopted sol-utions consist of low-strength mortarplugs or dewatered compartments infront of the shafts, enabling the tailvoid grouting around the segmentallining installed within the mined dia-phragm wall. The absence of a water-tight soil layer beneath the tunneldrive, together with the low strengthof the upper soil layers, made theseoptions very costly. Therefore, theprinciple of a bell inside the launchingshaft was chosen. Normally, a bell ismade by assembling reusable heavysteel parts at the beginning of eachtunnel drive, but as this assembly/dis-assembly procedure would be verycomplicated in the confined availablespaces, the choice was made to applya permanent concrete bell. At thearrival shaft, the TBM will arrive inan inundated shaft. After groutingthe void between the diaphragm wall

and tunnel lining, the shaft can bedewatered.

Design of the DiaphragmWalls

For the launching shaft, three differentsections were considered when design-ing the diaphragm walls, namely thefrontal wall, the strutted lateral walland the anchored lateral wall. For thelateral walls, the soil/water-retainingfunction is considered taking intoaccount a maximum soil/wall frictionand a zero soil/wall friction. This avail-able soil/wall friction depends on theamount of friction “consumed” forthe horizontal stability of the shaft, aswill be clarified in a following

paragraph. The frontal wall is designedwith a Young’s modulus of 12,000 MPa(cracked concrete/steel rebar) and5000 MPa (cracked concrete/GFRPrebar). Figure 3 shows the executionsequence and the envelope of the flex-ural moments for the lateral walls aswell as the maximum bendingmoments for all diaphragm walls andenvisaged scenarios. First, anchorpiles are installed from the existingsurface level in such way that thesteel anchor bar will protrude a fewdecimeters above the future exca-vation level. After a dry excavation(by means of dewatering inside the cof-ferdam) down to the level of −3.80 mNAP, the installation of the struttingframe and retaining anchors can takeplace. The latter are installed at alevel of −3.00 m NAP. Then, the dewa-tering is stopped, the water level insidethe pit rises up to a level of −1.45 mNAP and the wet excavation down tothe final excavation depth of−20.00 m NAP is performed. A 1 mthick underwater concrete slab ispoured after installing the steelanchor plates on the protruding steelanchor bars. Once the underwater con-crete slab has sufficiently hardened, thepit is pumped dry and the 1 m thickreinforced concrete base slab ispoured. Further on, the concreteworks of the permanent bell and inter-mediate floor slab take place. At thatmoment, the strutting frame can beremoved. To simulate the final stage

Fig. 2: Aerial view of launching shaft (section 1) and cofferdams for cut-and-cover part(sections 2 and 3)

Fig. 3: Cross-section of launching shaft, execution sequence and resulting bending moments

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of the structure, a neutral earthpressure (K0), instead of active/passive (Ka/Kp) earth pressures, is con-sidered, as well as a deterioration ofthe underwater concrete slab (loss ofsupport function). In the model of thelateral diaphragm walls behind the per-manent bell, the installation of theintermediate floor slab at −13.50 mNAP is modeled as well.

The increase in moment due to theloads of the bell on the floor slab isonly marginal (1.1%), while theabsence of shaft friction (δ = 0) leadsto an increase of 25%. The momentsin the anchored wall are reduced by17% compared to the strutted wall,owing to slightly less deep excavationlevel and the flexible support at thelevel −3.00 m NAP. The maximumdeflection of the diaphragm walls[service limit state (SLS) load case]depends strongly on the assumedwall friction angle. For a maximum δvalue (= 2/3 of the internal frictionangle ϕ value), the deflection islimited to 89 mm, while the maximumdeflection in case of δ = 0° mounts upto 147 mm.

Once the frontal wall is being minedthrough, its static system changes com-pletely. The remaining part above theTBM’s invert is coupled with thetympan wall and permanent bell, andtherefore no important sectionalforces remain inside the diaphragmwalls. The conservative approachwas chosen to load the part below theTBM’s invert by the bendingmoment in the lateral walls at asimilar height. This moment is intro-duced in the model with an oppositesign (Fig. 4).

The diaphragm panels are excavatedusing a 2800 mm wide grab. At thecorners of the construction pit, T-shaped panels are used. As the T-panels are primary installed panels,their width equals 2800 mm. Therequired width and length of the shaftresulted in panel widths between 6.4and 7.07 m comprising two rebar cagesof 2.85–3.00 m. At the interfacebetween the diaphragm walls and steelsheet pile walls, a U-normal profile(UNP) was attached to the diaphragmrebar cage and cast in (Fig. 5). Atypical sheet pile claw was weldedinside this UNP. To avoid concrete cov-ering up this claw, a temporary protec-tion plate was slid into the UNP andafterwards removed by vibrationalenergy before sheet pile wall installation.

Design of the Soft Eye

The models using a Young’s modulusof 5000 MPa for the GFRP-reinforceddiaphragm walls show a 5% lowerbending moment at the span, whilethe bending moment at the supportincreases by 21%. The support loadsare almost identical. As only a heightof 12 m consists of GFRP rebar andlow reinforcement stresses are encoun-tered, it was assumed allowable tomodel the entire frontal wall in thedetailed design stage with an E =12,000 MPa. For the deflections, thelower E value was relevant. An E =12,000 MPa showed a deflection of73 mm while in the case of an E =5000 MPa this deflection increased to148 mm. Adopting steel rebarØ50 mm, the required spacingbetween (vertical) bars could beachieved for the lateral diaphragmwalls. The GFRP rebar comprisesoval bars 40/75 mm or an equivalentdiameter Ø = 56.4 mm. Stirrups couldbe made as fully closed rectangles. Inthis way, spacious rebar cages wereconstructed (Fig. 6).

GFRP material has a linear stress–strain behavior until rupture occurs.However, the high-tension strengthtogether with a low E-modulus of theGFRP leads the concrete compression

zone to fail first. This means brittlefailure without any warning. There-fore, the resisting flexural capacity isdecreased to 80% according to designcodes2–4 and the concrete strength ofthe frontal wall is increased to C35/45(compared to C30/37 for the lateralwalls). The oval bars are producedwith a final coating of sand toimprove adherence. This adherencewas assumed to have an intermediatevalue between ribbed and non-profiledrebar, and lap and anchorage lengthswere defined on this basis. On site,pull-out tests were performed to vali-date this assumption. In contrast to tra-ditional diaphragm rebar cages, wherecage stability during installation isensured by welded rebar frames, thiswas not possible for the GFRP cages.Therefore, stability was achieved byimplementing two temporary steelrebar frames at the sides of the cages,which were removed bit by bit duringthe lowering of the cage.

Differential Behavior of theDiaphragm Wall Panels

Every diaphragm wall section isdesigned as an individual verticalbeam. Owing to the large spans andhigh loads, large deformationsresulted. A significant differential

Fig. 4: Situation when the tunnel boring machine has mined through the frontal diaphragmwall, imposed bending moment lower part of diaphragm wall. GFRP: glass fiber reinforcedpolymer

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behavior can be expected between theT-shaped panels that are aligned withthe lateral walls and the adjacentpanels of the frontal wall. The jointsbetween different panels are unrein-forced tongue-and-groove joints,opposing any differential deformationuntil rupture of a key occurs. Ques-tions were raised regarding the water-tightness in case such ruptureoccurred. Similar concerns aroseregarding the unreinforced sectionbetween both rebar cages of the T-shaped panels where lateral soil andwater pressures could shear off thelateral wall. For all panel joints, theresisting moment capacity dependson the acting horizontal compressiveload (i.e. in the plane of the panel).

This horizontal compressive loadvaries between 200 and 1500 kN/m.For this range, the ultimate bendingmoment was defined on the basis ofa triangular compressive stress distri-bution (Fig. 7). The resulting eccentri-city emax was defined by an overallsafety factor of 1.5 to obtain thedesign value of the allowable eccentri-city emax;d. For SLSs, the minimumcompressed concrete height was600 mm, thus avoiding tensile stressesat the location of the water barrier.For resistance to shear loads, theshear capacity of the key, neglectingthe contribution of the axial com-pression, is added to the shear resist-ance of the joint. The latter iscomposed of the contribution of the

adherence and friction over the com-pressed concrete section.

All consecutive execution steps areintroduced in a three-dimensionalfinite element calculation model. Thejoints between the diaphragm wallpanels have a rotational stiffnessaccording to Janssen’s theory, depend-ing on the acting bending moment/membrane force (M/N) ratios. Aninitial rotational stiffness, varyingalong the depth of the diaphragmwall, is implemented and the model isrun. The obtained M/N ratios areused to adapt the rotational stiffnesses,and after a few iterations, a final resultis achieved. The calculation showedthat possible ongoing cracking andleakage could be expected at theunreinforced zone between the rebarcages of the T-shaped panels, and thatshear forces could lead to rupture ofthe key in the joints between T-shaped panels and frontal wall dia-phragm panels. Both scenarios wouldhappen after dewatering the shaft, butonce the tympan wall had beeninstalled the risk would have gone.The contractor decided to place workbarriers at the corners for the safetyof workers in case of failure of thekeys. Furthermore, injection ventswere placed outside the diaphragmwall at the locations at risk to enableimmediate injection in case of encoun-tered leakages. In the end, none ofthese anticipated measures wasnecessary.

Global Stability of the Shafts

The horizontal stability of the shafts isa complex issue when they are situ-ated in weak soil layers combinedwith high hydraulic heads. Not onlydo considerable earth and water

Fig. 5: Diaphragm wall panel configuration and detail of connection with sheet piles

Fig. 6: Glass fiber reinforced polymer reinforcement cages using oval bars and closed stirrups

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pressures, which may be increased bythe temporary embankments, have tobe considered, but also the forcesexerted by the TBM have to betaken into account. The horizontalloads acting on the frontal wall canbe balanced by four resisting forces.These are the soil friction on thelateral diaphragm walls (Ffr;dw), thesoil friction on the temporary lateralsheet pile walls of the adjacent cut-and-cover tunnel sections (Ffr;shp),the passive earth pressure acting onthe compartment screen between the

launching shaft and cut-and-covertunnel (Fpe;cs), and the earth pressureacting on the end faces of the lateraldiaphragm walls (Fpe;dw) (Fig. 8). Allavailable resisting components wereintroduced in the calculation model,having a horizontal support stiffnessaccording to their individual load–deformation behavior. The SLSloads were first introduced into thecalculation model. Almost all resist-ing capacity was found by mobilizingthe soil friction on the diaphragmwalls. When increasing the horizontal

loads by a load factor of 1.5–2 and 3,respectively, one could see that thesoil friction on the sheet pile wallsand finally the passive earth pressureson the compartment screens and dia-phragm walls were used to a greaterextent. But even in the case of theload factor = 3, the encountereddeformations remained within accep-table limits (25 mm). The anchorpiles were assumed not to contributeto this horizontal resistance.However, the found displacementswere introduced as imposed defor-mations on the pile heads in the piledesign.

As the friction on the diaphragm wallswas fully mobilized, even in the SLSscenario, this friction could not beused a second time when designingthe diaphragm walls as retainingwalls. Furthermore, initial calculationresults made it clear that the separatediaphragm wall panels would behavelike books on a shelf when subjectedto these horizontal loads and thuslead to unacceptable deformations,

Fig. 7: Calculation model of the diaphragm wall panel joint

Fig. 8: Global stability of shaft, loads and resisting horizontal forces

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especially with respect to the under-water concrete floor, which maycrack. Therefore, it was decided toinstall a capping beam on top of thediaphragm wall panels, linking themtogether. This beam caused hightensile forces in the panels, requiring21 rebars, Ø50 mm and 12 m inlength, in every panel.

Tympan Wall and PermanentBell

Inside the launching shaft, a 1.00 mthick tympan wall is designed, situatedbetween the frontal diaphragm walland the permanent bell (Fig. 9). Theupper part of the diaphragm wall,twice mined through by the TBM, isconsidered as a temporary structureand therefore only designed for ulti-mate limit states (ULSs). In contrast,the tympan wall is designed for a100 year lifetime. The permanent bellhas only temporary functionality and isdesigned for the ULS only. Thetympan wall is designed as if nosupport from the permanent bellexists. The load encountered by theTBM’s cutting wheel is conservativelyassumed to be distributed over merely25 m soil friction on the alreadyinstalled tunnel lining. This means that8.8% (= 2.2 m tympan and frontal wallthickness/25 m) of the total thrustforce of the TBM is acting as a uni-formly distributed load along the per-imeter of the tympan wall opening.

To avoid the diaphragm wall panelsmoving forward while being minedthrough by the TBM, glued rebarsØ25 mm every 500 × 500 mm connectthe diaphragm wall with the tympanwall. This connection is intensifiedalong the perimeter of the openinginto Ø25–150 mm. To cope with thetail void injection pressures, the perma-nent bell has a circular reinforcementof Ø25–150 mm in the upper halvesand Ø20–150 mm in the lower halves,as the lower half is supported by thefloor slab and subgrade. This is alsothe reason why only the upper halvesneed shear reinforcement. The frictioncoming from the TBM thrust force andtransferred by the tail void injection istaken by a longitudinal reinforcementof Ø16–150 mm along the perimeterof the opening (Fig. 10).

The permanent bell, together with thetympan wall, is constructed in two con-secutive pours, whereby the formworkis used twice (Fig. 11). Fig. 10: Reinforcement schedule of the permanent bell structure

Fig. 9: Cross-section of launching shaft comprising soft eye, tympan wall and permanent bell.GFRP: glass fiber reinforced polymer

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Conclusion

In bored tunnel projects, the focus isvery often on the mechanized tunnelpart. Because all disciplines cometogether at the shafts and a widevariety of load cases acts upon them,these shafts are often the most challen-ging parts. Only an integratedapproach, where design, work prep-aration and executional aspects aretaken into account, will lead to theoptimum solution. The small spacingbetween the two tunnel tunnel boresin the Rijnlandroute project, togetherwith the high costs of traditional

starting plugs, led to the decision toadopt a permanent bell solution. Con-sideration of the global horizontal stab-ility of the shafts resulted in cappingbeams being placed on top of thelateral diaphragm walls. To ensureoptimum concreting, oval bars andclosed stirrups were adopted in theGFRP reinforced soft eyes.

Disclosure statement

No potential conflict of interest wasreported by the authors.

References

[1] Mortier H, Brugman M, Peerdeman B,Schubert T. The Rijnlandroute bored tunnel –continuously improving the mechanized tunnel-ing process. Proceedings ITA-World TunnelCongress Naples 2019.

[2] Guide for the Design and Construction ofConcrete Structures Reinforced with Fiber-Reinforced Polymer Bars -CNR-DT 203/2006.

[3] Beoordelingsrichtlijn glasvezelwapeningvoor toepassing als wapening in beton –

BRL0513 (in Dutch).

[4] Guide for the Design and Construction ofStructural Concrete Reinforced with Fiber-Reinforced Polymer (FRP)Bars –ACI 440.1R-15.

Fig. 11: Execution sequence tympan wall and permanent bell (top), installation of lower half formwork (bottom left) and lowering of thetunnel boring machine parts (bottom right)

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