External Post-Tensioning Retrofit and High- and Low-Cycle Fatigue Study of Connections and Joints in Steel-Concrete Box-Girder Bridges

Oreste S. BURSI Professor University of Trento Trento, Italy oreste.bursi@ing.unitn.it

Armando MAMMINO Civil Engineer S.I.GE.S. s.a.s Povegliano (Tv), Italy amammin@tin.it

Manuel FORTE Civil Engineer S.I.GE.S. s.a.s Povegliano (Tv), Italy manuel.forte1@tin.it

Nicola TONDINI Ph.D. student University of Trento Trento, Italy nicola.tondini@ing.unitn.it

Enrico TONON Civil Engineer S.I.GE.S. s.a.s Povegliano (Tv), Italy ing.enricotonon@libero.it

Summary This paper reports the study of a viaduct retrofit by means of external post-tensioning. The structure is a steel-concrete box-girder bridge and the retrofit has been necessary mainly owing to an advanced corrosion state which has affected the structural steel of the boxes. During the design two solutions have been conceived: i) cables with a parabolic development; ii) cables with a rectilinear development. The advantages and the drawbacks of these solutions designed by S.I.GE.S s.a.s. are here presented. In parallel with this problem, an on going study is being carried out on the: i) dynamic regime in order to identify the sensitive details; ii) validation of the Finite Element (FE) model by means of output-only ambient vibration tests; iii) optimization and physical testing of the critical details. This study is devoted to improve both the high- and low-cycle fatigue strength in steel-concrete box-girder bridges and the results achieved so far are here reported.

Keywords: steel-concrete composite bridge, external post-tensioning retrofit, high- and low-cycle fatigue behaviour.

1. Introduction Steel-concrete composite bridges represent a design option that is being increasingly adopted: i) in road networks; ii) in areas prone to high-intensity seismic events. The success of this design solution is due both to the advantages that composite elements offer in terms of stiffness, resistance and ductility and to the speed and ease of their erection. Moreover, the structural service-life has assumed a big importance in the design of structures, underlining the fact that the durability concept should be considered in the choice both of the material and of the structural typology. The durability of structural systems can be defined as the capacity to preserve the initial performance. It is thus linked to the capacity of the material of construction to keep its physical and mechanical properties unchanged within a given environment and under given working conditions. The steel-concrete composite action is particularly advantageous for bridges as it leads to enhanced stiffness, ultimate capacities and ductility. On the other hand, their non-homogeneity leads to significant problems at service conditions. Furthermore, it has to be considered that, for bridge structures, environmental conditions and types of loads increase the degradation of the material and in particular the most common degradation phenomenon is related to the corrosion of rebars in the concrete slabs. Bridges are, in fact, subjected to fatigue which leads to an increase of the cracks which, in turn, give rise to the penetration of the aggressive agents (e.g. chlorides, de-icing salts). In addition, an inaccurate maintenance, especially of the expansion joints, could lead to aggressive environments in closed box-girder composite bridges which could undermine the structural steel owing to corrosion. In this framework, an interesting technique for the rehabilitation and the retrofit of bridges is the post-tensioning by means of external slipping cables which can be straight or draped along the steel beam by means of deviators. Moreover, a peculiar issue is the lack of an accurate knowledge about dynamic effects, in particular in end-diaphragms and bearing regions which result to be very sensitive parts owing to stress

concentrations. In fact, they can be widely damaged when subjected to strong dynamic loadings, like earthquakes; in addition, dynamic effects induce vibrations owing to traffic loads, which can amplify fatigue phenomena. The paper is organised as follows. In Section 2 the analysis of the actual state of the bridge is presented, while in the Section 3 the FE model along with the dynamic analysis is shown; the localization of the sensitive details is also described. In Section 4 the on going validation of the FE model by means of dynamic identification tests is described, while in Section 5 the retrofit of the bridge by means of external post-tensioning with two solutions and different numerical analyses is presented. In Section 6 the optimization of the critical details and the forthcoming high- and low-cycle fatigue tests are described. Finally, in Section 7 the conclusions along with the future perspectives are drawn.

2. Analysis of the actual state of the viaduct The Montevideo or Vela viaduct is a steel-concrete box-girder bridge, built in the early eighties, with 7 spans, 75 m each. It is simply supported and 4 spans are curved with different curvature radius: going along the bridge from east to west the first three spans have R=350 m, while the 7th and last span has approximately R=200 m. The transversal section, see Fig. 1, consists of 4 lanes corresponding to 2 carriageways and 2 footpaths with a total width of the deck of 18 m. The total

height of the steel box is 4.10 m with a concrete slab of 30 cm. Moreover, the steel box is composed of transversal truss diaphragms which have different stiffness on the basis of their position along the span. The steel grade is S355 and the characteristic compressive strength of the concrete is Rck=40 MPa. The shear connection system between the steel box and the concrete slab consists of studs and T connectors.

The limited height of the viaduct from the ground permitted the positioning of propping structures during the construction stages. This allowed an amount of saved structural steel of about 20%. Therefore, in the analysis of the structure the various stages were taken into account in order to obtain the actual stresses in the bridge [1]. The necessity of the bridge retrofit by means of an external post-tensioning after only 25 years from its building is determined by essentially 3 factors: i) the traffic loads given by the national rules have risen in the past years; ii) the thickness of the steel plates which form the box is decreased by approximately 1 mm owing to a widespread corrosion process; iii) the traffic on the bridge cannot be interrupted owing to its strategic importance in the road network. The analysis of the bridge retrofit has been performed by the S.I.GE.S. s.a.s..

In fact, the midspan stress of a rectilinear span at the actual state with decreased thicknesses of the steel plates by applying the loads of the in force rules [2] is shown in Table 1. It is possible to observe that the stress in the steel is much higher than the yield strength fy=355 MPa.

3. FEM dynamic analysis of the bridge and identification of the sensitive details

3.1 Introduction The FE 3D model was developed by using the SAP2000 program [3]. As the viaduct consists of

4100

1000 16000 1000

60003000 3000

3000 6000 6000 3000

300

Fig.1 Transversal section at the bearing. Dimensions in mm

Table 1 Stresses at the bottom of the steel box and in the slab at midspan

MIDSPAN SECTION

DESCRIPTION MAXIMUM

STRESS AT THE BOTTOM [MPa]

MAXIMUM STRESS IN THE

SLAB [MPa] Maximum bending

and torsion 421.4 -15.5

simply supported spans, the analysis focussed mainly on two spans, in particular a straight span and the one with the smallest curvature radius, i.e. R=200 m. The dynamic analysis of the structure was carried out by means of a modal response spectrum analysis which allowed to identify the sensitive details.

3.2 Features of the FE model With regard to the model features, the bridge was discretized with shell elements for modelling the slab and the steel box, while frame elements were used for the truss elements of the transversal diaphragms and of the horizontal bracings. The composite interaction between the concrete slab and the steel box was made effective by body constraints which guarantee a full interaction as established by the European [4] and Italian rules [5]. The FE model of the curved span is shown in Fig. 2.

3.3 Dynamic analysis The bridge was dynamically analysed according to the national seismic code OPCM 3431 [6] which incorporates the EC-8 design philosophy. It is appropriate to underline that the case study is located in a zone of low seismicity, in particular Zone 3 according to the OPCM 3431. Hence, the bridge was seismically analysed and checked by employing a modal response spectrum analysis with ag=0.15g. The seismic analysis leads to a satisfactory behaviour of the viaduct in terms of check according to CNR 10011/88 [7] and CNR 10016/00 [5]; and it has been useful to identify the sensitive details from a dynamic point of view. As expected, the analysis showed that the bearing zone and the end-diaphragm were

locally the most stressed ones. In particular, the diagonals and the connections were identified as stress concentration zones to be particularly sensitive; Fig. 3 shows the longitudinal stress under the seismic load. Therefore, they could become critical in areas prone to high-intensity seismic events if not properly designed. Moreover, also the fatigue behaviour according to the part 2 of Eurocode 1, Eurocode 2, Eurocode 3 and Eurocode 4 was checked and it resulted satisfactory.

4. The validation of the FE model In the previous section, we identified the sensitive details of the bridge starting from a FE model which did not match the actual state of the bridge in terms of mass and stiffness owing to an advanced corrosion of the steel in the box and uncertainties related to time-dependent phenomena. Therefore, before choosing and optimizing the critical details we performed output-only ambient vibration tests in order to calibrate and validate the FE model. The use of output-only techniques is very useful when dealing with bridges, because they can be used without interrupting the traffic flow which is the main vibration source along with the possible presence of

Fig. 2 3D FE model of the curved span with R=200 m

Fig. 3 Longitudinal stress at the end-diaphragm of the curved span under the seismic load combination with ag=0.15g

Acquisition system position

Reference accelerometer

Transversal or longitudinal accelerometer

Vertical accelerometer

East

9.28 10.07 9.67 9.72 9.75 9.62 9.759.88

West

Fig. 4 Configuration A. Abutment side

wind. The test was performed with an acquisition system which can acquire till 16 signals and two types of accelerometers with different sensitivities were employed. Two accelerometer configurations as depicted in Figs. 4 and 5 were conceived for acquiring the most significant modes: they are represented by flexural, torsional and longitudinal modes of the box girder. The optimal accelerometer locations were selected by means of the AutoMAC matrices [8].

5. Retrofit of the bridge by means of external post-tensioning

5.1 External prestressing and post-tensioning in steel and steel-concrete bridges The concept of prestressing and post-tensioning is well-known, above all in concrete structures [9]. However, in steel-concrete structures, even though it can show good results, its application is very seldom and it assumes a slightly different meaning in the sense that we talk about prestressing when the external force is applied only to the steel part before the slab cast while with post-tensioning we mean the external force applied to the whole composite structure [10],[11]. Advantages arise from the prestressing because it is possible to obtain the same benefits with less applied force but it is suitable for new rather than existing structures. For the reason of dealing with an existing bridge, post-tensioning will be provided for. Moreover, since this solution is always external, an adequate study of restraint structures and deviators of forces as well as of possible local instabilities, which could lead to whole collapse of the structure, must be provided. Another important aspect is the choice of an appropriate development of the cables along the span. As a result, in this case two solutions were taken into account: i) a parabolic solution; ii) a rectilinear solution.

5.2 Solution with parabolic cables In a simply supported beam a parabolic development in the vertical plane of the cables is the most rational as the sections with higher bending moment are characterized by the maximum cable eccentricity. In fact, the cable is able to produce an equivalent load which is very close to an uplift uniform distributed load. Moreover, in correspondence to the bearings the post-tensioning force causes a shear decrease. In this case the cables were deviated also in the horizontal plane in order to avoid interferences with other structures like diaphragms, stiffeners etc. As a result, the cable development became quite complex as it is possible to observe in Figs. 6 and 7.

The post-tensioning force was calculated by considering the most stressed point at the bottom of the steel box at the serviceability limit state. The linear static analysis provided a value σ0=270 MPa. Hence, intending to fix σ0 to 190 MPa, the post-tensioning force became N=20496 kN and considering an amount of post-tensioning losses of about 12% it finally read N=22955.5 kN which meant 18 cables composed by 9 strands of 0.6’’ each for the

straight spans while 20 cables for the span with radius R=200 m. Although this solution presented many advantages, there were several drawbacks: the cable restraint structures were very complex for limiting the deformations and the stress caused by a high concentrated post-tensioning force Fig. 8.

Reference accelerometer

Transversal or longitudinal accelerometer

Vertical accelerometer

Acquisition system position

East

West

Accelerometer position on the pier

9.28 10.07 9.67 9.72 9.75 9.62 9.759.88

Fig. 5 Configuration B. Pier side

Fig. 6 Cable planimetry with parabolic solution

Fig. 7 Cable altimetry with parabolic solution

This entailed the use of plates with high thickness, till 30 mm, which determined difficulties in the organization of the building site, especially in terms of realization times and assemblage inside the steel box. With regard to the static analysis it was carried out with a simplified FE model composed of frame elements and the results in terms of stresses in the rectilinear span are summarized in Table 2.

5.3 Solution with rectilinear cables

In order to overcome the problems risen from the solution with parabolic cables, a second approach was investigated. It consists of a group of rectilinear cables lying on the bottom of the steel box. In this way the eccentricity of the cable with respect to the centre of the sections is higher and the losses owing to the angular deviations are practically negligible.

However, the benefit of the parabolic cable in terms of shear at the bearings is lost. Moreover in this case, in order to get a post-tensioning force proportional to the applied loads the cable anchorages were distributed along the span as depicted in Fig. 9.

The positions of the anchorages were conceived in order to obtain a stepwise constant bending moment diagram very close to the one obtained by the parabolic solution with the same number of cables, see Figs. 10 and 11. This positioning allowed a gradual increase of the post-tensioning force from the end to the midspan. Moreover, the critical anchorage received the thrust of only 6 out of 18 cables, this meaning restraint structures less complex and lighter. In fact, there was a steel reduction of 20÷25 percent owing to savings with respect to the steel used to anchor the parabolic cables.

Fig. 8 Particular of a cable restraint structure

Table 2 Parabolic solution. stresses at the bottom of the steel box and in the slab at midspan

MIDSPAN SECTION

DESCRIPTION MAXIMUM STRESS AT THE BOTTOM [MPa]

MAXIMUM STRESS IN THE SLAB [MPa]

Maximum bending and

torsion at ULS 345.7 -15.3

Maximum bending and

torsion at SLS 203.4 -11.0

Fig. 9 Cable planimetry with rectilinear solution

Fig. 10 Bending moment, shear and axial force due to the parabolic post-tensioning only

Fig. 11 Bending moment, shear and axial force due to the rectilinear post-tensioning only

Every anchorage consists of a rectangular box (80 cm × 35 cm and variable plate thickness 15 ÷ 25 mm) welded to the bottom of the steel box. In Fig. 12 is reported a typical restraint structure. The little eccentricity between the cable and the bottom of the box entails a local bending moment which determined a strengthening of the box as depicted in Fig. 13. Finally, this solution had the evident advantage of simplicity so that a standardization of the single parts could be achieved which turned out in a reduction of realization times and assemblage procedures.

5.4 FEM analysis by means of 3D models and comparisons

Besides the simplified FE model created with only frame elements, also two 3D models were developed in order to compare the results and to investigate the local effects in the restraint structures. Therefore, the static analysis of the retrofitted bridge was carried out: i) with the addition of parabolic cables; ii) with the addition of straight cables at the bottom of the steel box. In both cases the cables in steel box were modelled by means of frame elements and the post-tensioning force in the cables was given by applying a temperature gradient.

5.4.1 Static analysis of the solution with parabolic cables The static analysis was conducted by allowing for the construction stages in order to obtain the actual stress on the structure. Moreover, dealing with post-tensioned cables two different load conditions were considered: i) the post-tension of the cables at t=0; ii) the post-tension of the cables at t=∞, i.e. when the post-tensioning losses, such as friction, are completely occurred. The restraint structures of

the cables and the deviators were modelled with frame elements and they were connected to the cables by body constraints. In Fig. 14 the 3D FE model of the bridge with the parabolic solution is depicted. The compared results between the beam model and the 3D model of the curved span are reported in Table 3. The stress at the bottom of the steel box in each construction stage, but without the traffic loads, is shown and it is possible to observe that the stresses are very close to each other so to indicate a good agreement between the two FE models.

Fig. 12 Anchorage in correspondence of a transversal diaphragm

Fig. 13 Section A-A of the anchorage

Fig. 14 3D FE model of the curved span with parabolic cables Table 3 Stress comparison at midspan of the curved span

Stage σid,st,bottom [MPa] ∆ Beam model 3D model

1 2.8 4.5 -1.7 2 7.9 4 3.9 3 0.6 0.35 0.25 4 130 140 -10 5 56.5 50 6.5 6 1.2 2 -0.8

7.1 (post-tensioning t=0)

-112 -115 -3

8.2-1 0.1 0.5 -0.4 8.3 0.5 0.2 0.3 8.4 1 1 0 8.5 20.3 20 0.3 9.1 2 1.6 0.4 9.2 2.4 1.6 0.8 σ tot 225.3 225.75 MPa

5.4.2 Static analysis of the solution with rectilinear cables Besides the numerical analysis of the parabolic solution, also the 3D FE model of the solution with

rectilinear cables was developed as illustrated in Fig. 15. In this case the comparison between the parabolic and the rectilinear solution is reported in Table 4. The differences between the two solutions are highlighted only for the post-tensioning force. It is possible to observe that the rectilinear and the parabolic solutions lead to similar results, but with higher stresses for the solution with straight cables owing to the higher eccentricity.

6. On going activities related to the optimization and to the testing of the critical details

6.1 Employment of high-strength steel and welds Some details have been designed and others are still under study in order to improve the box durability. For instance, the anchorages of the cables in correspondence of the diaphragms are welded to the bottom of the steel box with only longitudinal welds. Therefore, the absence of transversal welds reduces the possibility of fatigue crack formation. In addition, to reduce the number of bolts, longitudinal welds are being introduced to replace bolt rows in connections and gusset plates. This solution is justified by the fact that it shows a better high-cycle fatigue behaviour at 2 millions of cycles than a bolted connection does [12]. See, for instance, the detail in Fig. 16 where the use of longitudinal welds will

increase detail category from 50 to 125. Moreover, longitudinal welded solutions will allow a reduction of costs and speed of erection. The employment of high-strength steel, such as S460, S500 or S550 [13], is also under study for these details in order to reduce weight.

6.2 Experimental tests by using dynamic substructuring Once identified the critical details of the bridge as the bearing regions and the end-diaphragms, the University of Trento is conceiving experimental tests to be performed in the Laboratory for Materials and Structures by using dynamic substructuring. The aim of these tests is to perform real-time high- and low-cycle tests on the optimized critical details by simulating the remaining life of the box girder numerically.

7. Conclusions This paper has shown the effectiveness of the external post-tensioning technique for retrofitting a steel-concrete box-girder bridge. Due to the complexity of the steel box and of the necessary restraint structures and anchorages the solution with straight cables lying on the bottom of the box resulted to be more competitive with respect to the one with parabolic cables owing to: less complicated restraint structures, shorter realization times due to reduced assemblage procedures. Moreover, bearing regions and end-diaphragms have been identified as sensitive steel connection

Fig. 15 Rectilinear span with straight cables at the bottom of the steel box. Table 4 Stress comparison at midspan between the parabolic and rectilinear solution due to the post-tensioning only

σid,st,bottom [MPa] Parabolic Rectilinear

t = 0 -90 -103 t = ∞ -86 -98

Fig. 16 End-diaphragm. Optimized welded detail

details for fatigue performance by the dynamic analysis. For validating these findings output-only ambient vibration tests were carried out in order to make possible the calibration of the FE model. Finally, the use of welded details as well as the employment of high-strength steel in order to increase the performance of the connections and to reduce the costs are under study.

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1990, pp. 93-98. [2] CONSIGLIO SUPERIORE LL. PP., Testo Unico - Norme Tecniche per le Costruzioni, Roma,

2005. [3] WILSON E. L., HABIBULLAH A., Sap 90 - A Series of Computer Program for the Static

and Dynamic Finite Element Analysis of Structures, Users Manual. Computer & Structures Inc., Berkeley, 1988.

[4] CEN, prEN 1994-2 (Final Draft). Eurocode 4: Design of Composite Steel and Concrete Structures - Part 2: General Rules and Rules for Bridges, Brussels, 2005.

[5] CNR, CNR 10016/00. Strutture Composte di Acciaio e Calcestruzzo: Istruzioni per l'Impiego nelle Costruzioni”, Roma, 2000.

[6] Ordinanza P.C.M. 3431, Norme Tecniche per il Progetto Sismico dei Ponti, Roma, 2005. [7] CNR, CNR 10011/88. Costruzioni di acciaio: Istruzioni per il Calcolo, l’Esecuzione, il

Collaudo e la Manutenzione. Roma, 1988. [8] EWINS D.J., Modal Testing 2nd Edition, Research Studies Press Ltd, Baldock, 2000. [9] POZZO E., Teoria e tecnica delle strutture. Volume III. Il cemento armato precompresso,

Pitagora Editrice, Bologna, 1999. [10] NUNZIATA V., Strutture in Acciaio Precompresso, Dario Flaccovio, Palermo, 2004. [11] TROITSKY M.S., Prestressed Steel Bridges. Theory and Design, Van Nostrand Reinhold

Company, New York,1990. [12] CEN, prEN 1993-1-9 (Final Draft). Eurocode 3: Design of Steel Structures - Part 1.9:

Fatigue, Brussels, 2004. [13] CEN, prEN 10025-6 (Final Draft). Hot rolled products of structural steels - Part 6: Technical

delivery conditions for flat products of high yield strength structural steels in the quenched and tempered condition, Brussels, 2003.