optimization of in situ construction of concrete decks: flexure tests of compact splices of...

Construction and Building Materials 41 (2013) 191–203

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Construction and Building Materials

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Optimization of in situ construction of concrete decks: Flexure testsof compact splices of reinforcement between phases

José Ramón Díaz de Terán a,⇑, José Turmo b, Juan José Jorquera c, Bryan E. Barragán d, Gonzalo Ramos c,Ángel Carlos Aparicio c

a Escuela de Ingenieros de Caminos, Canales y Puertos de Ciudad Real, Universidad de Castilla-La Mancha, Spainb Departamento de Ingeniería de la Construcción, Universidad Politècnica de Catalunya, Spainc Escuela de Ingeniería de Caminos, Canales y Puertos y de Ingeniería de Minas, Universidad Politécnica de Cartagena, Spaind BASF Construction Chemicals Europe, Italy

h i g h l i g h t s

" Flexure tests of loop joints between phases have been performed." SLS and ULS have been studied." Conventional formulae are valid to predict the ULS and modifications have been introduced to predict crack width." Results support the use of loop joints between different concrete phases and provide accurate data to predict.

a r t i c l e i n f o

Article history:Received 31 January 2012Received in revised form 4 October 2012Accepted 21 November 2012Available online 8 January 2013

Keywords:Concrete viaductsSpan-by-span castingMovable Scaffolding SystemCompact splicesCritical path methodLoop joints

0950-0618/$ - see front matter � 2012 Elsevier Ltd. Ahttp://dx.doi.org/10.1016/j.conbuildmat.2012.11.085

⇑ Corresponding author.E-mail address: [email protected] (J.R. Dí

a b s t r a c t

This article presents a study of an alternative casting process for concrete bridge decks executedspan-by-span through self-supporting launching falsework or Movable Scaffolding System, focusing ona reduction of the critical path of construction. The efficiency of non-standardized reinforcement splicesis evaluated through pure flexure static tests. Noteworthy, due to its geometrically compact characteris-tics, these splices do not interfere with the movement of the internal formwork of the deck. These originaltests demonstrate that such splice solution for connecting reinforcement represents a safe and econom-ical stress transference alternative for different types of concrete classes and performances.

� 2012 Elsevier Ltd. All rights reserved.

1. Introduction

Besides precast beams and the incremental launching method,there is another classical construction procedure for built concretedecks for long viaducts with medium spans; the span-by-spanmethod with Movable Scaffolding System (MSS). However, evenif the method is fairly industrialized, its competitiveness is reduceddue to its demanding execution time.

This work discusses a modification of the construction processof bridges executed span-by-span with MSS, focusing on the reduc-tion of the critical path. At the same time, the paper analyzes thestructural aspects involved in such improvement.

ll rights reserved.

az de Terán).

Self-supporting launching falsework are being used since thesixties. The first time this system of construction was used, wasin Germany. The Krahnember viaduct, designed by Hans Wittfoth,was built in 1961 [1]. It was mainly from the seventies when thespread across Europe took place. Some of the most noteworthyperformances of that time are the Glattfelden Lättenbrücke via-ducts, Ponts sur le Viaduc du telent Chavornay, and Lac de la Gru-yere, in Switzerland.

The usual spans achieved by the MSS method are in the range of40–60 m. The traditional incremental launching procedure consistsin the execution of the inferior slab and flanges of the cross sectionin a first phase (Fig. 1a), subsequently executing the upper slab(Fig. 1b). Then, once the necessary concrete strength for prestress-ing is achieved, the tendons are stressed and the falsework ad-vances to the next span. This sequence generally requires2 weeks per span, although this period can be reduced by reducing

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Fig. 1. Traditional method of construction by phases. (a) Phase 1 and (b) phase 2.

192 J.R. Díaz de Terán / Construction and Building Materials 41 (2013) 191–203

curing times and continuing reinforcement and splicing activitiesat night, thus involving much higher costs. In both, the traditionalincremental procedure and MSS, the falsework advances span byspan, setting up the casting joint at a distance equal to 0.2 L fromthe piles, where L is the length of each span, so that the bendingmoments at the joint between longitudinal phases are as low aspossible.

MSS represents a great advantage from the point of view ofmodern requirements of Occupational Health and Safety [2,3] sinceinvolves an industrially prefabricated auxiliary structure that per-mits the use of a platform on which the collective security mea-sures are implemented at the factory. Hence, operating risks arelower than for other systems.

In terms of reduction of execution time [2] and, therefore, con-struction costs, it has been observed that a further evolution of theMSS method is based in building partial self-supporting schemesthat allow to move the falsework forward, and later complete thetransversal section by simpler auxiliary means, out of the criticalpath. Then, in order to improve the performance and executiontime of each span of the deck, a construction sequence by transver-sal phases, is proposed. This sequence is different from the tradi-tional one and a priori allows executing one span per week, withactivities out of the critical path. This variation makes it a muchmore competitive methodology.

The new method (Fig. 2) considers the execution of the U-beamand flanges in a first phase (i.e. bottom slab, webs, and flanges ofthe upper slab). The first stage of prestressing takes place oncethe elements have reached the required strength. The prestressingforce introduced in the structure is generally in the range of 50–60% of the final value. This introduces self-supportability to the exe-cuted span, so the falsework can move forward to the next span.The second phase, corresponding to the central area of the upperslab, can be executed later without disturbing the movement ofthe falsework, i.e. outside the critical path and with simpler auxil-iary means.

There are a series of viaducts already executed in Spain follow-ing this construction sequence [4–6]. These examples have proventhe following advantages with respect to the previous procedure:

– Permits the generation of a resistant core that accelerates theadvance of MSS.

– Allows the use of simpler auxiliary means to execute the secondphase, out of the critical path.

– Improves the construction performance and timing; from 1.5 toless than 1 week per span.

– Avoids the aesthetic issues at the web-flange union area.– No uncovered prestressing ducts in phase 1.

Fig. 2. New method of construction by p

– From the construction point of view, clarifies the load distribu-tion between the falsework and the deck, since the falseworkonly has to support the weight of the first casting phase. Theweight of the second casting phase is carried by the self-supporting core.

– Introduces an increase of capacity of the existing falsework,since it does not have to support the weight of the whole sec-tion but only that of the first casting phase, until prestressing.

However, the application of this procedure is constrained byconnecting reinforcement and splicing issues linked to the castingphases. Specifically, the problem arises when connecting the rein-forcement of the slab of the second phase with that of the firstphase. Respecting the classic conditions for reinforcement splicingby overlapping provided by Codes and Standards [7–12] makes thesolution unviable in practice. Regulatory requirements oblige toleave long protruding rebars to connect the two phases. In practice,this fact complicates the extraction of the interior formwork of thesection (Fig. 3). On the other hand, the use of mechanical splices,that would also ensure a compact geometry, would invalidatethe solution due to its high cost. The use of welding to splice rein-forcement on site is forbidden in Spain and in many neighboringcountries due to the difficult quality control and loss of ductilityof the welded reinforcement. Transversal prestressing of the slabis also a possible solution that would allow connecting the phases,but it is not always an economical option and does not have uni-versal acceptance.

Consequently, for the successful implementation of the newmethod, it would be necessary to develop a compact splice geom-etry that would allow the removal of the inner formwork of thedeck in an industrialized way. In this way, the straight protrudingrebars (Fig. 4a) that constitute a layer of reinforcement impedingthe extraction of the inner formwork, are replaced by loop-typesplices or loop-joints (Fig. 4b) that to do not represent an obstaclefor the removal of the internal formwork. This article presents anexperimental study that supports the use of this type of splicegeometry, that is not covered by current Codes and Standards.Tests evaluate the fulfillment of the performances required by cur-rent regulations both in Ultimate Limit State (ULS) and Serviceabil-ity Limit State (SLS), under pure static flexural loading. This is donefor a given splice geometry in slabs cast in phases with differenttypes of concrete; normal strength concrete, self-compacting con-crete, and high strength concrete.

The behavior of slabs with loop joint under dynamic loads wasstudied at the Laboratory Luis Agulló of Structural Technology, atthe Polytechnic University of Cataluña, Spain [13]. A total of 8 slabswere tested under dynamic simple flexural bending. All the tests

hases. (a) Phase 1 and (b) phase 2.

Fig. 3. Inner formwork for the first phase.

(a)

(b)Fig. 4. The two types of connections in the cross section: (a) straight connectionand (b) loop-joint connection.

J.R. Díaz de Terán et al. / Construction and Building Materials 41 (2013) 191–203 193

were carried out until the slabs cracked. Two different types of loopjoint were adopted: the first one, with 20 mm bars 20 and the sec-ond one with 25 mm bars. The slab dimensions were5.60 � 1.60 � 0.285 m. The results of the dynamic tests show thatunder cyclic loading the stiffness gradually reduces. The level ofstresses that were obtained under a dynamic loading of2,000,000 cycles were beyond the values that the different norma-tive sets, so the compact splices (loop joints) can be used under dy-namic loading and fatigue conditions.

2. Flexural behavior of splice with compact geometry

There are several Codes that propose formulae for crack openingcalculation (Model Code, Spanish Standard EHE 2008, Eurocode 2)that were taken into account for this study. These formulae do notconsider the effect of loop joints or the existence of a casting joint:

Model Code proposes:

Wk ¼ ls;maxðesm � ecm � ecsÞ ð1Þ

where Wk is the crack width, ls,max is the length of the bar where rel-ative displacement between steel and concrete takes place, esm isthe medium strain of the steel, ecm is the medium strain of the con-crete, ecs is the concrete strain due to shrinkage.

Were as Spanish Standard EHE 2008 proposes: Ac/As

Wk ¼ bSM eM ð2Þ

SM ¼ 2c þ 0:2sþ 0:4k1UAc=As ð3Þ

where Wk is the crack width, b is the coefficient that relations thecrack opening and the characteristic value, SM is the average spacebetween cracks, eM is the average strain of the reinforcement, c isthe concrete cover of tensile reinforcement, s is the space betweenbars, k1 is the coefficient that takes into account the influence of thetensile diagram, Ac is the concrete area at the tensile zone, As totalarea of tensile reinforcement, and U is de diameter of reinforcementbars.

Eurocode 2 proposes:

Wk ¼ b SMeM ð4Þ

SM ¼ 50þ 0:25 k1k2U=qT ð5Þ

where Wk is the crack width, b is the coefficient that relates theaverage crack opening and its characteristic value, SM is the mediumspace between cracks, eM is the average strain of the reinforcement,k1 is the bonding coefficient that depends on the type of bar, k2 is acoefficient that takes into account the shape of the strain distribu-tion, qT = As/Ac, and U is de diameter of the bars of reinforcement.

There are technological applications that use compact splices ofloop-joint type, as the inverted T-beam system (also called PoutreDalle). This system allows eliminating the formwork and providesa safe working area, and it is also efficient in terms of cost and tim-ing. An improved version of this method was used to execute theCenter City Bridge, Houston County, Minnesota Bridge 6679 [14].However, current Codes do not include loop-joint compact splices[15], i.e. the connection of two slabs through a double closed loopreinforced with transversal rebars, even if references and casestudies on the subject date back to the seventies [16].

There are bibliographic references dealing with the behavior ofloop-joints under tensile loading. The behavior is explained fromthe following theories, among others:

– Theory of bond between steel and concrete, assuming a stressdistribution or anchorage length [11,8,17].

– Radial force method [18], where the design is done in functionof the tensile stress in the rebars, the rebar curvature and diam-eter, and the distance between adjacent loop-joints. Thismethod considers the placement of transversal rebars acrossthe loop, and is applicable to cases where bending takes place.

– Bond and radial force mixed method [19] in which the impor-tance of the longitudinal splice is highlighted in order to reducecracking. Considers the addition of the force that is transferredthrough the bond between steel and concrete, and the force thatis transferred radially.

– Method of superposition of forces [20], that considers the forceabsorbed by the concrete, by the bond between steel and con-crete, and by the transversal reinforcement across the loop.

– Method of failure at the loop-joint plane [21].– ‘‘Free body’’ analysis of the cross section just before crack for-

mation through a single hook developed by C. Tan, S. Yeo andC. Hon between 1998 and 2001, at the National University ofSingapore [22].

– Strut and ties models developed between 2000 and 2005 at theNational University of Singapore by Ng and Ong [22].

Regarding the behavior of loop-joints under flexural loading, aseries of tests have already been proposed in order to study the sta-tic and dynamic response, between precast elements or a combina-tion of precast and in situ construction:

Fig. 5. Detail of loop joint connection.


– Tests carried out in Germany, at the Institut für Beton und Sta-hlbeton of Karlsruhe University [23]. Highlighted the influenceof the relative position of the loops and the overlap, the coverconcrete, the presence of transversal reinforcement, and thebar diameter.

– Studies at the Dutch TNO-IBBC (Institute TNO for BuildingMaterials and Building Structures) [24] defined a formulationfor the strength of the loop-joint as a function of the diameterand area of steel of the loop and the transversal reinforcement,concrete tensile strength, length of the splice, and the distancebetween the exterior loops and the external edge of the testedbeam. Results highlight the importance of the transversal rein-forcement to increase strength and control the crack openingmode, and the significance of leaving a sufficiently rough sur-face between phases.

– Tests performed by the Building Research Station (Israel Insti-tute of Technology) on slabs with double loop-joint connectionbetween two precast elements. It was found that the first cracksappeared at the casting joint, at relatively low load levels [25].

– Static and dynamic tests at the Korea Institute of ConstructionTechnology. Tests on combined sections consisting in a steelbeam and a concrete slab for the study of cracking and compat-ibility [26]. Also tests on precast slabs under 3 and 4 point bend-ing [27] in which instead of overlapping rebars, differentanchorage lengths of loop-joints were tested according to theACI Code and DIN 1045 Standard. The results of this experimen-tal program showed that the longer the anchorage length, thebetter the crack opening response.

– Tests on slabs with loop-joints at the National University of Sin-gapore [28,22,29] evaluated the different methods of analysis ofloop-joints descried before, as well as its finite element model-ing. This study included the less conventional loop-joint solu-tions, as the horizontal loop-joints [30].

– Tests at the Chulalongkorn University of Bangkok (Thailand), inwhich the results on slabs (without casting joint) where com-pared with foreseen results from finite element analysis [34].

Test results obtained in Korea confirmed that the behavior ofthe loop joints is similar to that of reinforced concrete beams withconventional reinforcement and depends up to a great extent onthe bar diameter and the disposition of transversal reinforcement[27]. This is in accordance with results by Charuchaimontri et al.[31]. In terms of crack widths, it was observed that the responseis better than that calculated from Standards if overlaps larger than350 mm are provided, while with 250 mm splices, service condi-tions were not satisfied [26,27]. The cracking load under flexuralmoments is lower than that obtained without casting joint. Thisdepends up to a great extent on the roughening technique imple-mented, and on the interface preparation before creating thein situ joint [24,25,29]. Also, it was observed that the initial crackis larger than that predicted by Codes. However, as the load in-creases, the values tend to converge, and finalize below predictedvalues; i.e. Eurocode 2 [26].

In the last years, a series of bridges have been built in Spain inwhich the traditional casting sequence has been modified [6,4,5].The lack of specific Standards applicable to this construction tech-nique and the consequent impossibility of a compact splicinggeometry, made necessary an experimental program to demon-strate the structural capacity of loop-type reinforcement detailswith no overlapping to assure the connectivity between castingphases.

This paper presents the results of the experimental programwhere the fulfillment of the performances required by current reg-ulations both in Ultimate Limit State (ULS) and Serviceability LimitState (SLS), under pure static flexural loading is tested (Section 1),for loop-type reinforcement splices, loop-joints (Fig. 5) in slabs

fabricated with different types of concrete, including casting joints,and subjected to normal stresses produced by pure static flexure.The experimental program was carried out at the Structures andMaterials Laboratory of the School of Civil Engineering of CiudadReal (University of Castilla-La Mancha).

Tests carried out simulate elements cast by phases, withoutprefabricated components. This means that the data presented ap-plies to the behavior of the joints between casting phases of a realviaduct constructed in situ. Hence, it does not apply to joints be-tween precast segments. This is why the tested joints presentasymmetric configurations [unlike joints between prefabricatedelements [26,27,24,30,22,25], and fabrication methods and form-work simulate those really used on site. Moreover, the work notonly presents the study of the capacity of the joint against normalstresses (ULS) like in [20,24,22,30], but special attention is given toservice conditions, particularly those that influence durability, likethe crack opening. Also, this work evaluates the extrapolation pos-sibilities from traditional simple formulae to predict the flexurecapacity and crack opening of these splicing geometries, and avoidsthe use of formulae that are certainly more accurate but at thesame time more complex. Such complex approaches require theknowledge of new scientific theories that are not always available,and are too far from the practice of designers and contractors. Onthe other hand, even if it is common to find research studies onthe structural behavior of bridges in SLS [32–38], or regardingstrengthening to extend its life span [39,40], it is much moreuncommon to find studies on the structural effects of the construc-tion process of medium-span bridges in general, and on bridgesbuilt up by MSS.

3. Tests

The main principal stresses to which the casting joint is subjected are thosecoming from the transversal flexure of the upper slab. This is why the tests per-formed in this study involve slabs under pure flexure. Results will be useful to studythe behavior under SLS and ULS.

3.1. Materials

Concrete components included a CEM I 52.5R cement, crushed limestone sand(0–4 mm) and gravels (5–12 and 12–20 mm), siliceous sand (0–4 mm), and twochemical admixtures from BASF Construction Chemicals, a policarboxilate-basedhigh range water reducing admixture (HRWRA, Glenium C-355), and a viscositymodifying agent (VMA, Rheomac 350). Mix designs provided to the precast planttargeted 35 MPa and 70 MPa for NSC and HSC, respectively. However, modificationsimplemented at the industrial plant, based on previous experience, knowledge ofthe machinery and production procedures, led finally to the mix proportions pre-sented in Table 1. Concretes were mixed in a 1 m3 pan mixer, during summer time.As it can be observed from Table 1, the NSC has a higher content of coarse aggregatethan HSC and SCC. This could increase the roughness of the joint in the case of NSC,and lead to a higher interlocking between faces.

Table 1Mix proportions (kg/m3).

NSC HSC SCC

Cement I-52.5 R 350.0 457.0 457.0Crushed limestone sand 0–4 mm 931.0 1050.0 1050.0Natural limestone sand 0–4 mm 161.0 220.0 220.0Crushed limestone gravel 5–12 mm 512.0 400.0 400.0Crushed limestone gravel 12–20 mm 300.0 100.0 100.0Total water 140.0 167.0 167.0HRWRA 5.5 3.2–4.0 9.0VMA 2.2

Table 3Compressive strength results of the different concretes and phases.

NSC fcm (MPa) Phase 1 Phase 2N-NSC-1/2 53.0 54.5N-NSC-C 61.5 56.8

HSC fcm (MPa) Phase 1 Phase 2N-HSC-1/2 69.2 55.4N-HSC-C 62.4 66.7

SCC fcm (MPa) Phase 1 Phase 2N-SCC-1/2 60.7 74.1N-SCC-C 72.2 69.9


Steel used for reinforcement was a B 500SD (UNE 36065 EX), with a limit of pro-portionality of 500 MPa, ultimate tensile strength of 575 MPa, elongation at break16%, and a total elongation at maximum load of 7.5%.

Workability of NSC and HSC measured by the slump test correspond to a S3consistency level (EN 206-1), NSC-1/2 and HSC-1/2 presented both a slump valueof 130 mm, while NSC-C and HSC-C presented slump values of 140 and 110 mm,respectively. Table 2 includes the fresh state measures of SCC, flowability, and pass-ing ability under unconfined and confined conditions, measured through slump-flow, J-ring, L-box, and V-funnel tests ([41–44], respectively).

Table 3 includes the average compressive strength, fcm, at 28 days [45–48], foreach concrete and slab. Values represent an average of 3 cylindrical specimens of150 � 300 mm. As it can be observed, the compressive strength of NSC is consider-ably higher than foreseen.

Fig. 6. Simulated position of the slabs within the bridge deck.

3.2. Large-scale specimens

The slabs consist of 0.285 � 0.60 � 2.90 rectangular prisms. The element simu-lates a piece of the slab of the bridge deck (Fig. 6), and it is cast in 2 phases. The aimwas to have a representative thickness with respect to usual dimensions for thesetypes of viaducts. The casting of the 2.90 m long slab was done in two phases, a firstphase of 1.25 m length and a second phase of 1.65 m.

The behavior of the joint under normal stresses (N) is evaluated for three typesof concretes; normal strength conventional concrete (NSC), self-compacting con-crete (SCC), and high strength concrete (HSC). Three tests have been carried outfor each material. The reference slab is called control slab (C), with continuous lon-gitudinal reinforcement along the entire length of the slab (Fig. 7). The other twoslabs (Fig. 8) comprise a reinforcement configuration that assures the connectivitybetween phases in a loop-type compact way. Table 4 presents the nomenclature ofthe 9 tests carried out.

The used loop-joints (see Fig. 5) consisted of 3 loops of 20 mm rebars in eachjoint. The loops were transversally connected by 6 rebars of 16 mm diameter. Theloops exceed the face of the first casting phase by 248 mm. The splice length ofthe loop of the second casting phase is also 248 mm. In the case of the control slabs,the transversal reinforcement consisted in 6 longitudinal rebars (3 at the top and 3at the bottom) of 20 mm diameter. Cover concrete was 30 mm in all cases.

The fabrication of the slabs took place at a precast concrete production plant inBarcelona, during the summer of 2008. Two metal molds were utilized, where theelements were cast horizontally, Figs. 9 and 10. To assure the positioning of thereinforcement, no welding was permitted. Even if welding is common practice inthe Spanish precast industry, welding is usually forbidden on site, and in this casewould have also generated misleading results.

Table 2Fresh state measures of SCC.

Slump-flow Measure

T50(sg) 1.31–5.19Df (mm) 610.00–685.00J-ring with 10 mm bars Ring 10DfJ (mm) 550.00–630.00H1(mm) 25.00–55.00H2 (mm) 15.00–35.00Df–DfJ (mm): 60.00–120.00

V-funnelTv (sg) 5.40–9.20L-Box with 3 U12 bars Bars U12T60 (seg) 1.88–6.03H1(mm) 100.00–155.00H2 (mm) 60.00–85.00Density (ton/m3): 2.38–2.41Temp. (�C) 28.00–30.00

The molds were sprayed with demolding agent before casting. The slabs werecast in two phases. For the definition of the joint, the casting of the first phasewas done against a bulkhead made out of a phenolic wood board. This bulkheadwas later the mold for the second pouring phase, i.e. the joint between castingphases. The second phase was executed 1 week after the first one, previously wet-ting the surface of the first phase joints to avoid the absorption of water from theconcrete of the second phase. Since there were only 2 molds, each concrete type in-cluded 4 batches; 2 for the control slabs N-C and 2 for slabs N-1 and N-2, that wereexecuted at the same time. NSC and HSC were consolidated by means of a pokervibrator. SCC slabs were not compacted at all. After casting the second phase, theentire element was wetted and covered with burlap fabric; such curing took placeduring 7 days.

As mentioned at the previous paragraph, the formwork of the joint betweenphases consisted in a phenolic wood board, leading to a very smooth interface sur-face. The joint did not receive any kind roughening treatment. The use of a nervo-metal bulkhead to increase the roughness and interlocking at the joint, was alsodiscarded. The aim was to work on the safe side, and validate the use of this typeof splices under whichever working hypothesis, without imposing any specifictreatment to the joint. Evidently, an increase of the surface roughness would in-crease the test results [29].

3.3. Test configuration and procedure

All slabs were subjected to a 4-point bending, see Fig. 11. Hence, the centralportion of the slab, and therefore the joint, remains under pure bending stresses.The load application is made by a servo-hydraulic actuator of 1000 kN capacityand 300 mm of piston displacement. The test is carried out under displacementcontrol. The two symmetrical loads were separated 0.60 m by means of a steelbeam supported on steel rollers. The slabs were simply supported, with a free spanof 2.4 m (see Fig. 11). Longitudinal displacement and rotation was allowed at bothsupports.

Data recorded along the test include the load applied by the actuator and mea-sured by its load cell, vertical displacements at the supports and casting joint, andthe horizontal displacement or crack opening between the faces of the casting joint(Fig. 11). As it can be seen in Fig. 12, vertical displacements at the supports are re-corded by means of a 10 mm gauge length LVDT placed in an auxiliary structure, atnorth and south supports. Also supported by auxiliary structures, two verticalLVDTs are placed at the joint, close to the laterals of the slab; sides east and west.Two additional LVDTs of 10 mm gauge length were attached horizontally to thebottom of the slab at sides east and west, to measure the opening of the castingjoint, Fig. 13. The calculation of the deflection and joint opening along the loadingphase is done by half adding the displacements measured by each pair of LVDTs,so that arithmetic mean is obtain, since no significant difference between pairswas found in the tests. A data acquisition system records the measures from alltransducers at regular time intervals.

Tests were performed approximately 6 months after the fabrication of the ele-ments. After callibration of the transducers, a pre-loading of approximately 30 kNwas applied, in order to settle and stabilize the entire test setup. Afterwards, the

Fig. 7. Detail of reinforcement of control slabs SCC-N-C, NSC-N-C y HSC-N-C (dimensions in mm).

Fig. 8. Detail of reinforcement in loop-joint slabs; SCC-N-1/2, NSC-N-1/2, HSC-N-1/2 (dimensions in mm).

Table 4Nomenclature of the elements.

Loop joint connections Straight connections

Normal strength concrete NSC 1/NSC 2 NSC CHigh strength concrete HSC 1/HSC 2 HSC CSelf compacting concrete SCC 1/SCC 2 SCC C


test starts by applying a constant displacement rate of 1.2 mm/min. Loading isstopped before the complete collapse. The observed crack pattern is similar in allcases; Fig. 14 shows the crack patterns in case of a control slab (Fig. 14a) and a slab

Fig. 9. Formwork and reinf

with loop joint (Fig. 14b). Even if several cracks appear between loading points, themajor and first crack to appear is the one in correspondence with the casting joint.The average distance between cracks is approx. 150 mm, see Fig. 14.

4. Theoretical predictions

The behavior in service, and strength of the joints was predictedby the formulae included in different codes [7,8,10,12]. This formu-lae estimate de behavior of the joint as a function of the reinforce-ment ratio effectively bonded in the cross section under study.

orcements of the slabs.

Fig. 10. Different concrete phases of the slabs.

Fig. 11. Simplified scheme of the flexural tests with the LVDTs (dimensions in mm).

Fig. 12. Test configuration.

Fig. 13. LVDTs to measure the crack opening between faces at the joint plane.


Hence, to predict the response in service and ultimate capacity, aperfect splicing of the reinforcement across the joint plane, ishypothesized. This also implies assuming that the used splicinggeometry is effective and permits the total stress transfer betweenreinforcements. This hypothesis is verified experimentally, as ex-plained below.

Regarding the capacity of the element at ULS, according to theCEB-FIP (1993), Eurocode 2 and EHE-08, the beam fails with a

deformation plane in domain 2. Meaning that the strain of the steelreinforcement in tension is 10‰ and that of the upper fiber be-tween 0‰ and 3.5‰. The equilibrium equations are set accord-ingly, and the validity of the initial hypothesis is subsequentlyverified.

On the other hand, for the calculation at ULS through ACI-318-05, a plastic design has been considered, assuming the unitarydeformations in concrete are directly proportional to the distanceof the neutral axis, and the maximum unitary deformation of theextreme compression fiber is equal to 0.003.

In all cases, steel was considered to have a perfect elastic–plas-tic response. A rectangular stress diagram was used to account forthe contribution of concrete in ULS. No safety coefficient forstrength reduction or long-term actions, were adopted.

In this way, the ultimate moment, Mu, values at the joint planeand the load, Q, that produces that moment, are obtained for thedifferent slabs as a function of the 28 day concrete compressivestrength, see Table 7.

Regarding the cracking behavior in service, both EHE-08 andEurocode 2 have similar approaches; both start from the evalua-tion of the critical moment of cracking, Mcrit (when the crackstarts). However, the 2 codes differ in the value of the coefficientsthat take into account the effects on the crack opening displace-ment of the different physical phenomenons involved. The valuesof Mcrit for the tested elements, are presented in Table 5.

The cracking moment at the cross section of the joint is pre-sented in Table 5, for each type of concrete. The calculation of thesevalues was done by considering the mean tensile strength of theconcrete, estimated from the characteristic compressive strength,according to Eq. (1) [12]. The Model Code CEB-FIP (1993), section

(a) Control slab (b) Slab with loop-joint

Fig. 14. Cracks in the element during the test.

Table 5Comparison between theoretical and experimental values of the cracking moments.

Mcrit tests Mcrit theoretical c (Mcrit tests/Mcrit theoretical)

NSC 1 9.11 31.19 0.29NSC 2 9.61 31.19 0.31NSC C 9.88 31.19 0.32HSC 1 10.24 32.32 0.32HSC 2 5.47 32.32 0.17HSC C 8.98 32.32 0.28SCC 1 7.27 34.19 0.21SCC 2 4.79 34.19 0.14SCC C 10.91 34.19 0.32


7.4.3.1, proposes different formulae to estimate the crack openingvalue as a function of the actuating load.

fctm ¼ 0:30f 2=3c=k sifck � 50 Mpa

fctm ¼ 0:58 f 1=2ck si f ck > 50 Mpa

ð6Þ

Eq. (6) mean tensile strength fctm.

5. Results, analysis and discussion

Cracking formulation depends on cracking moment. This crack-ing moment can be theoretically calculated from the tensilestrength of the concrete. However tests show that the cracking mo-ment when a joint is present is considerably lower than the theo-retical cracking moment and it do not depends on the type of splicethat is at the joint (Table 5). That is why a correction factor isneeded. This correction factor must be included in the formulaeto obtain the crack opening. This enables the use of the equationsfrom the standards with a very accurate prediction of crackopening.

The average value of the ratio between the experimental andtheoretical cracking moments is ~amed ¼ 0:26, with a typical devia-tion of 0.07, and a characteristic value ~acharact ¼ 0:14 (assuming anormal distribution of results). The characteristic value will hereaf-ter be used as a correction factor for the theoretical cracking mo-ment of the different Codes, in order to study the validity of theanalytical estimations.

Load–deflection curves are presented in Figs. 15, 17 and 19, forNSC, HSC, and SCC, respectively. The slight settlement produced atthe supports has been subtracted from deflection values at the castjoint plane, hence a net deflection is considered. The load- crackopening responses of the same elements are presented in Figs. 16,18 and 20. A net crack opening has been considered by subtractingthe concrete elastic deformation to the measure of the horizontalLVDTs at the bottom of the elements. Tables 6 includes the actuat-ing moments at the joint, producing different crack openings ineach case.

5.1. Normal strength concrete

The three NSC slabs tested present a very similar behavior interms of vertical deflection, stiffness and ductility, as it can be ob-served from Fig. 15. The crack opening response is also very simi-lar, at least up to a crack opening of 0.3 mm. After this crackopening, the control slab seems less sensitive to the increase ofthe actuating moment. In all slabs, the first cracks appeared atthe casting joint. Note that the real cracking moment (y axis valuein Fig. 16) is lower than that theoretically calculated according toEHE-08 (Table 5); the real cracking moment is around 0.14 timesthe calculated one. As it can be observed from Table 6, the theoret-

ical predictions are considerably conservative for all evaluated de-sign codes (CEB-FIP 1993, EHE-08, and Eurocode 2). At all levels ofcrack opening, the real experimental moment is always lower thanthe one predicted by the codes, which leaves the design at the un-safe side. This effect can be graphically observed from Fig. 16,where the theoretical curve (wk EHE-08) that relates the actuatingmoment to the crack opening is always above the experimentalresponse. Taking into account that the displacement does not de-pend on the splice geometry, the difference of behavior is justifiedby the presence of the casting joint. Considering studies by [29],this phenomenon is maximized by the way in which the castingjoint was actually executed; i.e. with phenolic wood panels thatgenerate a very smooth surface, instead of other techniques thatproduce higher roughness.

Since the real cracking moment of the slabs is significantly low-er than the theoretical value calculated by EHE-08 (0.14 Mcrit the-oretical, y-axis value in Fig. 16), the theoretical formulationproposed by this code was modified by applying a coefficient equalto 0.14 to Mcrit. Including this value in the EHE-08 formulae, ahighly accurate fitting is obtained with the experimental values.Hence, the modified formulation represents a validated criteria toestimate crack openings in this type of joints.

The same conclusions can be drawn from the analysis of resultspresented in Table 6, which summarizes the experimental valuesand those from modified code formulations, for different crackopenings according to different exposition classes. The originaltheoretical predictions according to all codes are on the unsafe sidefor all elements and crack opening values. If the suggested correc-tion (modified code predictions with correction factor) is consid-ered, the degree of precision for the estimation of the crackopening is very high.

In any case, considering that the differences of behavior be-tween the control and loop-joint slabs are minimum within thecrack opening range that matters in terms of SLS design of struc-tural elements (up to 0.4 mm), it can be assumed that the influenceof the type of splice does not affect to a great extent the response ofthe joint in terms of cracking. It seems that it is rather the actual

0 10 20 30 40 50 600

50

100

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350

0

50

100

150

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250

300

350Q

(kN

)

DEFLECTION (mm)

N-NSC-1

N-NSC-2

N-NSC-C

Fig. 15. NSC; load–vertical deflection response.

0.0 0.1 0.2 0.3 0.4 0.5 0.60

20

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100

120

140

0

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wk EHE modified

MO

MEN

T (k

Nm

)

CRACK WIDTH (mm)

wk N-NSC-1

wk N-NSC-2

wk N-NSC-C

wk EHE

Fig. 16. NSC; experimental moment Mk-crack width wk responses and theoreticalpredictions.

0.0 0.1 0.2 0.3 0.4 0.5 0.60

20

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140

0

20

40

60

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wk EHE modified

MO

MEN

T (k

Nm

)

CRACK WIDTH (mm)

wk EHE

wk N-HSC-1

wk N-HSC-2

wk N-HSC-C

Fig. 18. HSC; experimental moment Mk-crack width wk curves and theoreticalpredictions.

0.0 0.1 0.2 0.3 0.4 0.5 0.60

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140

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20

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140

wk EHE modified

MO

MEN

T (k

Nm

)

CRACK WIDTH (mm)

wk N-SCC-1

wk N-SCC-2

wk N-SCC-Cwk EHE

Fig. 20. SCC; experimental moment Mk-crack width wk curves and theoreticalpredictions.

0 10 20 30 40 50 600

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100

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350

0

50

100

150

200

250

300

350

Q (k

N)

DEFLECTION (mm)

N-HSC-1

N-HSC-2

N-HSC-C

Fig. 17. HSC; load–vertical deflection curve.

0 10 20 30 40 50 600

50

100

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0

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300

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Q (k

N)

DEFLECTION (mm)

N-SCC-1

N-SCC-2

N-SCC-C

Fig. 19. Vertical load–deflection curve (SCC).


existence of the discontinuous plane of the joint and possibly itslack of roughness that principally affects cracking [24,25,29] Cur-rent codes provide a formulae that does not consider the fact ofthe presence of a casting or cold joint. Thus, it is logical the under-estimation of the crack opening for a given bending moment. Thisis why a modification of the formulae taking into account the realmean tensile strength value at the joint plane (correction factor), isproposed.

5.2. High strength concrete

From the load–deflection curves of the HSC slabs (Fig. 17), avery similar response from control, and N-HSC-1 and N-HSC-2,can be observed. All slabs present a high ductility, even when slabN-HSC-2 presents a higher deformability than its companion ele-

Table 6Comparison between experimental results and modified theoretical values (Mcrit modified).

Crack width (mm) Concrete Element Experimental Moment (kN m) EHE-08 modified (kN m) EC-2 modified (kN m) CEB-FIP 1993 modified (kN m)

0.1 NSC N-NSC-1 23.8 16.1 18.0 19.4N-NSC-2 24.9N-NSC-C 26.3

HSC N-HSC-1 23.9 16.2 18.1 19.5N-HSC-2 20.4N-HSC-C 24.9

SCC N-SCC-1 21.8 16.4 18.2 19.7N-SCC-2 21.4N-SCC-C 24.5










Table 7Comparison between experimental and theoretical values of failure loads.

SLABS Failure load (experimental) CEB-FIP 1993/EC-2 and EHE-08 ACI-318

Q (kN) M (kN m) Q (kN) M (kN m) Q (kN) M (kN m)

N-NSC-C 320.0 146.8 247.8 114.3 245.1 113.1N-NSC-1 300.0 137.8 246.1 113.6 243.3 112.3N-NSC-2 300.0 137.8 246.1 113.6 243.3 112.3N-HSC-C 280.0 128.8 249.9 115.3 247.3 114.1N-HSC-1 300.0 137.9 247.2 114.1 244.5 112.8N-HSC-2 290.0 133.3 247.2 114.1 244.5 112.8N-SCC-C 310.0 142.3 252.4 116.4 249.7 115.2N-SCC-1 310.0 142.3 249.3 115.0 246.7 113.8N-SCC-2 310.0 142.3 249.3 115.0 246.7 113.8


ments. Again, the type of splice does not seem to affect the overallcapacity of the element.

The crack opening response of HSC elements can be observedfrom Fig. 18 and Table 6. Again, four aspects are worth to be high-lighted, (a) the very similar response of the slabs at least up tocrack openings of approx. 0.3 mm, (b) the significantly lowerexperimental cracking moment with respect to formulae providedby considered codes, (c) the underestimation of the crack openingby the different formulations for a given bending moment, and (d)that a correction of the theoretical Mcrit value considering the realvalue obtained in the tests (0.14 correction factor) permits a highlyprecise estimation of the crack opening.

5.3. Self-compacting concrete

From the observation of the load–deflection and moment- crackwidth curves presented in Figs. 19 and 20, respectively, and thevalues summarized in Table 6, including the test results for SCC,a similar behavior can be observed for the three slabs. From theseresults, the following aspects can be highlighted, (a) the compara-ble stiffness, deformability, and capacity of the three slabs, (b) thesimilar SLS behavior for all slabs, at least up to crack openings ofapprx. 0.4 mm, (c) the lower experimental cracking moment, withrespect to Code predictions (test values are in range of 1/3 of thetheoretical value), (d) the underestimation of the crack opening


by the different formulae for a given moment, if Mcrit is not cor-rected by its real value (0.14 times the theoretical estimation).

5.4. Comparison among results of tests with different concretes

Table 7 summarizes the ultimate loads obtained experimen-tally, as well as predictions by existing formulae contained inCodes [7,8,10,12]. As it can be observed, the ultimate load of allslabs is very similar (practically equal). Also, failure loads do notshow clear trends, i.e. control elements, with continuous longitudi-nal reinforcement, do not necessarily show a higher capacity thanelements with loop type splices. On the other hand, the experimen-tal and predicted failure loads are independent from the concretetype and compressive strength. This can be explained by the factthat it is basically the reinforcement ratio that controls the failureof the joint. This shows that the ULS capacity is hardly sensitive tothe different levels of mean compressive strength of the concrete,but is determined by the strength of steel, that is the same in allcases. The experimental failure loads (ULS) are within a range of280–320 kN, in all cases. Also, the predictions by codes and recom-mendations are systematically on the safe side, meaning that theuse of the traditional formulae for the prediction of the failure ofthe joints can be used without modifications (Table 7). Note thatthis validation of the current formulae has been done under pureflexure, with no shear or axial stresses involved. In any case, itseems reasonable to assume that the existence of shear stressesdoes not invalidate the direct use of the formulation. In contrast,the possible presence of axial tensile stresses at the joint is verylimited, of relatively small magnitude, and only present in the caseof transversal sections with inclined webs. Only in the case of crosssections with strongly inclined webs (hence, with important axialtensile stresses), the use of the current formulations would bequestionable. Finally, the ductile performance of the slabs supportsthe fact that the loop joint does not introduce any kind of brittlebehavior in the slab. Thus, involving no safety reduction with re-spect to other type of connecting reinforcement.

The influence of the lack of roughness at the joint due to the useof phenolic wood could have determined the initiation of the crackopening at that section, between the two construction phases ofthe slabs. Table 6 summarizes the theoretical critical moments ofcracking according to Codes, and the experimental values from testresults. As it can be observed, the real value of the critical momentof cracking is in the order of 0.14 of the theoretical Mcrit. Since thisresponse is independent from the type connecting reinforcementactually used, the result is attributed to the existence of a verysmooth casting joint. In any case, the crack opening behaviorwould be notably higher in presence of rougher joint conditions.However, considering the lack of experimental results and workingon the safe side, the method developed in this article can beadopted to evaluate the cracking behavior of the joints.

As it can be observed for the three types of concrete, a very goodapproximation of the real values can be achieved by substituting inthe formulae the theoretical value of Mcrit by the real value ob-tained from tests (0.14 times de theoretical value: correction fac-tor). In any case, given the equal behavior of the slabs with loopjoints and those with continuous reinforcement, it is reasonableto assume that this issue is not related with the type of connectingreinforcement but with the actual existence of the discontinuityplane of the joint. Hence, the application of a correction factorequal to 0.14 is proposed when phenolic wood joints are used;the worst situation that can take place in practice. Any treatmentor construction method increasing the joint interlocking wouldalso permit an increase the value of such coefficient.

It must be noted that the flexural moments and axial stressesactuating on the joint when in service, are very low. To calculatethe opening of the joint when in SLS, these values correspond to

the combination of semi-permanent actions in reinforced concrete.The transversal flexure moments come mainly from traffic loads,primarily through local flexure of the upper deck, and to a lesserextent from the own weight and permanent loads. Axial stressesare generally very low with respect to the flexural moment.

Given the safety coefficients used in ULS for dominant loadsð~a ¼ 1:5Þ, and the load combination coefficients for the dominantsemi-permanent load combination (Ø2 = 0.2) [49,50], and assum-ing that the actuating moments come primarily from the variableloads, a cross section designed according to the theoretical ulti-mate moments included in Table 7 would have a maximum servicemoment (in case of strict ULS design) of approx. of 13% of the trans-versal flexure design moments (112–117 kN m). With the studiedconnection, the service moment would be in the range of 14.5–15.2 kN m, which represents a crack opening smaller than0.1 mm in the tests (the maximum crack opening measured inthe experimental program was 0.075 mm for a 15.2 kN m, in testN-HSC-2).

Also, it should be noted that the different concrete strength ormix design, i.e. higher content of fine materials in the case ofHSC and SCC, has not significantly influenced the behavior of thejoint, neither influenced the critical moment of cracking. Thebehavior of the three types of concrete was very similar.

6. Conclusions

This work presents an experimental study aiming to evaluatethe response of different types of splices as connecting reinforce-ment in the case of reinforced concrete elements cast in phases.The considered splices include out-of-Norm loop-joint compactsplices. Control elements with normalized configurations whereused as a reference. The casting joints where subjected to normalstatic loads. Three different types of concretes were considered;normal strength, high strength, and self-compacting concrete. Thisstudy aims supporting the modification of the construction pro-cesses of concrete bridges build by the span-by-span methodthrough self-supporting launching falsework. Since the main nor-mal stress actuating on the casting joints is that coming from thetransversal flexure of the upper deck, the tests performed consistin longitudinal slabs tested under pure flexure. Test results permit-ted the study of the behavior of the joints at SLS (cracking) and ULS.

Firstly, as a key conclusion of this paper, test results have pro-ven the adequacy of the loop-joint splice geometry under studyto resist normal static stresses at SLS and ULS. At the same time,the use of a compact splice geometry would allow demoldingand earlier movement of the interior formwork of the deck, makingthe modified construction method even more competitive with re-spect to the current solution.

The comparison between crack openings measured in the testedslabs permits concluding that the behavior of the joint is indepen-dent of the type of connecting reinforcement, as it can be assumedfrom the similarity of the cracking moment results obtained foreach splicing condition, for the three slabs in each case. However,the cracking prediction provided by current Codes is in the unsafeside in all cases, particularly for low stress levels. The direct use ofthe formulae provided by Codes to estimate the cracking momentin presence of the type of joint reinforcement studied in this work,is not recommended. The cracking at low load levels is attributedto the existence of the discontinuity joint with smooth faces, whichinvolves an important reduction of the critical moment of crackingwith respect to the theoretical value. In view of the fact that thecharacteristic critical moment of cracking in the tests is 0.14 timesthe theoretical value, the incorporation of this reduction factor inthe theoretical formulae, is proposed. In this way a high degreeof precision is achieved when in the crack opening range of 0.1–


0.4 mm. In any case, despite the fact that the presence of the jointdecreases the cracking strength, in real bridges stresses involved atSLS are low with respect to ULS. Hence, the cracking induced at SLSwould be little and would not compromise durability. From a prac-tical point of view, and focusing on the improvement of the behav-ior in service, the execution of a joint as rough as possible isproposed, using nervometal or brushing the surface after the cast-ing of the first phase. [24,25,29].

In terms of ULS, test results indicate failure loads in the range of280–320 kN, which are higher than the estimations from the for-mulae of Codes analyzed in this study, in the range of 243–253 kN. This results permits concluding that Codes can be usedwithout modifications for ULS design of loop-joint connecting rein-forcement, for the three types of concretes considered in this study(NSC, HSC, and SCC). On the other hand, the use of this type ofsplices does not seem to introduce a reduction of structuralductility.

Therefore, based on the obtained test results and theoreticalestimations analyzed in this work, it can be stated that loop-jointconnecting reinforcement is applicable with the same safety mar-gins than continuous and straight splices to transfer static flexurestresses, using conventional design procedures included in currentCodes.

Acknowledgements

The authors would like to thank the collaboration of the techni-cal department of Pacadar, and the personnel of their precast plantlocated in Sant Boi de Llobregat (Barcelona, Spain), for the fabrica-tion of the slabs. The tests presented in this article were carried outat the Materials and Structures Laboratory of the School of CivilEngineering of Ciudad Real, of the University of Castilla – LaMancha (Spain). The authors would like to thank the personnelof the laboratory and its director, Dr. Xiaoxing Zhang, for thesupport and professionalism. At the same time, the significantcollaboration of Civil Engineer Santiago Salinas Clavero duringthe experimental program, is greatly appreciated.

The study presented in this paper is part of the doctoral re-search of the first author of the paper. It was done as a coordinatedproject between the Technical University of Catalonia (UPC) andthe University of Castilla – La Mancha (UCLM), funded by theSpanish Ministry of Science and Innovation. The global project(BIA2006-15471-C02-01) has been directed by Prof. GonzaloRamos (UPC), from an original idea of Prof. Angel Aparicio (UPC).This research was part of the sub-project BIA2006-15471-C02-02,directed by Prof. José Turmo. The funding from the Spanish Minis-try of Innovation and Science through the cited projects, for theexecution of the tests, is greatly appreciated. As well as the fundingfor the corresponding analysis from the Ministry of Science andInnovation of the Government of Castilla – La Mancha throughProjects BIA2009-13056 and PII2I09-0129-4085.

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[41] EN 12350-8:2010. Testing fresh concrete – Part 8: Self-compacting concrete –Slump-flow test.

[42] EN 12350-12:2010. Testing fresh concrete – Part 12: Self-compacting concrete– J-ring test.

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optimization of in situ construction of concrete decks: flexure tests of compact splices of...

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