files.transtutors.com€¦ · web viewfatigue crack behavior at rib-to-rib butt weld connection in...

Fatigue Crack Behavior at Rib-to-Rib Butt Weld Connection in Orthotropic Steel Bridge Deck

Saifaldien-shakir1, Fatih Alemdar1, 2

1 Yildiz Technical University, Faculty of Civil Engineering, Department of Civil Engineering, Structural Division, Davutpaşa Campus, 34210 Esenler, Istanbul, Turkey

Abstract

Rib-to-rib (RR) butt welded connections are one of the most sensitive locations for

encountering fatigue failure in orthotropic steel decks (OSDs), and numerous fatigue cracks

arising from these areas have been identified in existing OSD bridges. This study focuses on

the fatigue cracking process, fatigue characteristics, and failure mechanics of RR connections

under cyclic loading. Experiments were performed on six trapezoidal rib specimens, including

five specimens with initial cracks and one control specimen. Furthermore, one centric load

case was considered in order to investigate the ease of fatigue cracking occurrence at the

bottom of the existing longitudinal ribs, on the area between the connections plate and

channels, from the practical fatigue failure cases at the weld joints. Static loading was firstly

conducted with the aim of measuring the elastic strain distributions at the strain gage

readings, based on which critical positions were identified, and the strain deformation values

obtained from the strain gage locations applied to each work method were evaluated. High-

cycle repeated loading was subsequently implemented, from which the fatigue crack

propagation process, fatigue failure mode, characteristic fatigue life, and stiffness curves were

obtained. Three cracks growth stages were mainly observed, as determined by the dye-

penetrant and dynamic stiffness crack detection methods. Moreover, the fatigue propagation

life was plotted as a crack length for each specimen against a number of cycles, and the

failure cycles were recorded for the fatigue strength S–N Curves in AASHTO (2007).

1Corresponding Author: Tel: Fax: E-mail addresses:

1

Keywords: Orthotropic steel bridge deck; rib-to-rib butt weld connection; crack propagation; fatigue life; finite element analysis.

1. Introduction

Orthotropic steel decks (OSDs), composed of deck plates that are orthotropically stiffened by

longitudinal ribs and transverse diaphragms, have been extensively used in various steel

bridges owing to their benefits such as light weight, high load-bearing capacity, and expedient

construction, among others. In engineering practices over the past several decades, many

improvements have been achieved in all aspects of the design, fabrication, inspection, and

maintenance of such bridge decks, and the structural behaviors have subsequently been

significantly enhanced. However, fatigue resistance remains a prominent issue, and even a

design controlling aspect for OSDs, as these are directly and repeatedly loaded by vehicle

tires. An additional adverse effect concerning OSD fatigue behavior is that numerous welds

between steel plates causes the local stresses near the weld toes and roots, including welding-

induced residual stresses and live load-induced stress ranges, concentrate to high levels as a

result of structural discontinuities, which is commonly regarded as the major cause of material

fractures following numerous load cycles. Following the first report of OSD fatigue cracks in

the Severn Bridge (Britain) in 1971, a great deal of fatigue cracking damage has been

identified in actual OSD structures worldwide. Fatigue cracks have also been detected in

several OSD bridges in China. In total, over 1000 cracks have been observed in the Hu Meng

Bridge, Guangdong Province (1997) since their first appearance in 2003, and rehabilitation

technologies using high ductility performance concrete and the sandwich plate system have

been applied comparatively. More than 100 fatigue cracks were found during a regular

inspection of the Jiang Yin Bridge, Jiangsu Province (1997), which is a suspension bridge

with a steel box girder, with the first crack being detected in 2011. Furthermore, as illustrated

in Fig. 1, several fatigue cracks have been reported in the Fatih Sultan Mehmet Bridge in

2

Turkey. Fatigue cracks have been detected, and have gradually increased, in OSDs, generally

originating from repeated out-of-plane bending owing to repeated wheel loads: fatigue cracks

initiate at the welded joint between the connection plate and trough rib, and propagate to the

rib wall. Such fatigue damage may cause a sudden decrease in the longitudinal rib stiffness.

Out-of-plane bending results in high local flexural stresses in the welded joint between the

connection plate and trough rib, because the plate and trough rib thicknesses are relatively

small.

[Figure 1 about here]

Numerous efforts have been devoted to the fatigue resistance of OSDs since their first

appearance in bridges [1-3]. In OSDs, the most common rib-to-deck welded connections are

also the most fatigue-sensitive locations owing to direct wheel loading, and various

researchers have focused on this topic. Sim et al. [4] investigated the effects of both weld

melt-through and distortion control measures on the fatigue resistance of the rib-to-deck plate

welded joint, by employing six full-scale orthotropic deck specimens, and determined that no

crack was identified at the full penetration weld roots, while pre-cambering was beneficial to

fatigue resistance. As part of continuous work [5], these authors further performed finite-

element analyses of the test specimens using the effective notch stress method, and results

demonstrated that increasing the deck plate thickness was effective in reducing stresses, while

the rib wall thickness had little effect. Xiao et al. [6] carried out local stress analysis and

fatigue evaluation of rib-to-deck joints in OSDs under the action of wheel loads using finite

element (FE) software, where the surface stresses in the deck plate were revealed to be

significantly larger than those in the rib wall in the case of 75% weld penetration into the rib

wall. Ya et al. [7] reported on the fatigue tests of rib-to-deck specimens with three weld

details: 80% partial joint penetration (80% PJP), weld melt-through (WMT), and both. The

3

results indicated that specimens with WMT appeared to exhibit slightly lower fatigue

strengths than the 80% PJP specimens, but the difference is more likely to be within a usual

test data scatter, meaning that both details have comparable fatigue strengths. Liu et al. [8]

also constructed refined solid models using a multi-sub-model technique in order to analyze

the structural hotspot stresses in rib-deck welded joints, where the effects of the weld profile,

weld toe radius, and mesh size were discussed. Pfeil et al. [9] numerically investigated the

complex stress distribution and concentration at the rib-to-deck welded connection of

orthotropic decks with trapezoidal closed ribs under heavy wheel loading. In addition to rib-

to-deck connections, the rib-to-floor beam welded joints, cutouts in diaphragms, and other

welded connections have attracted the attention of researchers. Connor [10] systematically

analyzed the effect of cutout geometry on stress in the region of the welded rib-to-diaphragm

connections in OSDs, and proposed an improved geometry based on an FE parametric study.

By employing a full-scale laboratory test, Connor and Fisher [11] presented a procedure for

calculating stresses at the rib-to-diaphragm connection, which is consistent with the fatigue

resistance provided in the AASHTO (2012) LRFD bridge design specifications [12]. Choi and

Kim [13] evaluated the stress characteristics and fatigue crack behavior of longitudinal rib-to-

crossbeam joints, as well as the effectiveness of crack repair, using a stop-hole in an

orthotropic steel bridge. Based on analytical and experimental studies, Oh et al. [14] presented

optimal parameters regarding the height, thickness, and shear area of a cross-beam, and the

efficiency and shapes of bulkhead plates, in order to prevent fatigue cracks. Moreover, Xiao et

al. [15] experimentally and theoretically investigated the causes of fatigue cracks in the

longitudinal ribs of Kinuura Bridge (Japan), and concluded that a lack of penetration zones

resulted in low fatigue strength of the butt-welded joints. Numerous stress analyses, as well as

fatigue life estimations, have been conducted for rib-to-rib (RR) connections where fatigue

cracks have usually been observed near or in the butt-welded joints. Another experimental

4

and analytical works were also conducted by Yokozeki and Miki [16] to evaluate the fatigue

performance of connections between longitudinal U-ribs and transverse ribs with and without

the slit on the transverse rib web, to Clarify the relation of tire load locations to SHSS

working at the longitudinal-to-transverse rib connection, Confirm the applicability of the

SHSS approach to the connection of OSD , and Evaluate the fatigue life of the connections

with the slit and without the slit. Therefore, based on previous investigations, this study

focused on the sensitive connection details of fatigue cracking with the presence of initial

cracks at the bottom of the existing trapezoidal rib-to-rib butt welded connections from the

longitudinal direction through experimental and analytical investigation.

2. Material Properties

All of the plates used to fabricate the specimens were S235 steel material. Following the static

and dynamic calibration test, tensile coupons were cut from the web (rib steel plates) and

flange (deck steel plates) of specimen 2. The rib coupons (RC1 through RC3) had a 6 mm test

thickness, while the deck coupons (DC1 through DC3) had a 12 mm test thickness, and the

tests performed in accordance with ASTM E8/E8M (2009) [17]. The measured material

properties are presented in Table 1, which indicates the elasticity modulus, yield stress, and

ultimate stress. The test was conducted using an Instron machine, as illustrated in Fig. 2.

[Table 1 about here]


3. Analytical Investigation

3.1. Details of Longitudinal Ribs

The trapezoidal longitudinal rib examined analytically in this study is illustrated in Fig. 3. The

cross-sections of the rib studied in this case provide the details of existing ribs used in the

Fatih Sultan Mehmet Bridge in Istanbul [18].


5

3.2. FE Modeling of Ribs

In this study, the longitudinal rib and deck plate components used in orthotropic bridge decks,

as well as part of a plate under load, were modeled using the finite element program

ABAQUS [19]. The surface area of the plate under load increased when an axial load was

applied; thus, a larger surface area was accounted for in the FE analysis. The model width was

defined as the longitudinal rib depth.

The longitudinal ribs, upper plate, welds, and connection plate were defined using S235 steel

material. The Young’s modulus and Poisson ratio are given as 200 GPa and 0.3, respectively.

The plate under load was modeled using a rigid material, because the most important aspect

of the analysis was to evaluate the longitudinal rib and not have any unexpected deformations

in the plate under load. All of the materials in the FE model were defined using a solid

homogeneous material type. The boundary conditions of the longitudinal rib and upper plate

at each end location were constrained during the displacement and released during the

rotation.

The connections between the plate under load and upper plate, as well as the longitudinal rib

and plate, were modeled using a tie constraint in the analysis. Hard contact interaction and

friction were defined between the plate under load and upper plate.

The plate thicknesses were 6 mm and 12 mm for the longitudinal rib and upper plate,

respectively, while the weld thickness and length around the connection region were defined

as 8 mm and 5 mm, respectively. A 6 mm-thick connection plate was used to combine the

upper plate and longitudinal rib. The interaction among the connection plate, weld, and

longitudinal rib was selected as the tie constraint in order to provide the full connection. All

analyses performed were linear elastic, which means that yielding and other non-linear effects

were not of interest [20].

6

The mesh size for the entire model was approximately 25 mm, except for the connection

plate, welds, and small part of the ribs in the middle of the model. The mesh in these regions

was very fine sized, at 5 mm, with an eight-node linear brick element (C3D8R) defined for

the mesh element, as illustrated in Fig. 4. An axial load of 100 kN of concentrated force was

applied at the center of the plate under load, which is considered as the maximum load of one

wheel when a truck is fully loaded with a dynamic amplification factor.


3.3. RR Connection Side

In this study, the connection region of the longitudinal ribs was considered. This region is

generally constructed 1 m from one end of the diaphragm plate under the bridge deck in real

bridges [21]. An FE model of the connection plate and weld to be used to combine the upper

plate and longitudinal rib was created. The weld thickness was selected as 8 mm, as

mentioned previously, and applied to the connection plate thickness as a butt weld, as

illustrated in Fig. 5. Full interactions; that is, the tie constraints, between the plate and weld,

and between the weld and longitudinal rib, were provided.


3.4. Analytical Results and Discussion

The FE model was constructed as if the loading passed in the middle of the longitudinal rib.

The bottom path of the rib was subjected to maximum principal stresses and deformations

compared to the other system components. However, the weld stresses around the connection

plate were concentrated, and the connection region could be susceptible to increased fatigue

cracking, as illustrated in Fig. 6. It can be observed that fatigue cracks began to occur on the

7

closed side of the bottom of the rib. The stresses on the trapezoidal ribs exhibited single-sign

stress distribution, particularly the compression region along the force direction; that is, the

stresses on the trapezoidal rib were generally tension stresses. Moreover, the maximum

deformation locations were closer to the bottom side of the trapezoidal rib, as indicated in Fig.

7. This result explains where the cracks began to grow, as also illustrated in Fig. 1. These

findings are in line with actual situations.



The analytical approach in this study was also focused on determining the strain values in the

same strain gage positions that were fixed on the specimen in the static calibration test, in

order to facilitate simulation between the ABAQUS program and experimental test. The three

different longitudinal paths to be examined analytically are illustrated in Fig. 8. These paths

were taken along the bottom of the model, which considered the most critical locations and

would be exposed to the greatest deformations. Path one passed through strain gages 1 and 3,

path two passed through strain gage 2, and path three passed through strain gages 4 and 5.

The details for these gages are displayed in Table 2. The strain values E33 at each gage

location were taken with respect to the Z-axis, as illustrated in Fig. 13.



4. Experimental Investigation

4.1. Test Setup

4.1.1. Specimens

8

Six full-scale RR butt-welded connections were manufactured. Each specimen was composed

of a deck plate, two trapezoidal ribs, and connection plate, and all specimens had identical

geometrical dimensions. The deck plate thickness t d was 12 mm and the rib plate thickness t r

was 6 mm. The specimens were 400 mm wide in the transverse direction and 1500 mm long

in the longitudinal direction. All ribs had a height of 256 mm and top width of 310 mm. The

rib arc curvature radius was 25 mm. The rib legs were welded to the deck plate by means of

PJP, while complete joint penetration (CJP) was applied at the RR connection edges, which

were produced by carbon dioxide gas arc welding, in strict accordance with the American

National Standard Welding Code – Steel [22]. The geometries of the RR connection are

illustrated in Fig. 9.


4.1.2. Test Rig

Prior to testing, the longitudinal rib and deck plate were horizontally erected and fixed to the

foundation "fixture" using fastening holders. The RR specimens were simply supported at the

two ends of the deck plates and ribs. Each holder sat on the longitudinal fixture at the plate

(400 mm length, 150 mm width, and 50 mm thickness), which was welded to the holder and

fixed to the fixture by using rivets to prohibit possible transverse rigid body movement of the

specimen during the cyclic loading process, as illustrated in Fig. 10. The fixture was 1500 mm

long, 360 mm wide, and 240 mm high. The specimens were tested experimentally using two

methods: static calibration and dynamic fatigue tests. The work was conducted by an Instron

test machine, model 8803.


9

4.2. Test Procedure

4.2.1. Measurement of Strain Gages

Five strain gages were installed on the control specimen from the rib bottom side in the

longitudinal direction in order to measure the strains near the RR butt weld connections, as

illustrated in Fig. 11. The first and most important step in the strain gage installation process

was surface preparation, which was necessary to produce a chemically clean surface with

appropriate roughness, alkalinity, and layout lines. The surface preparation was performed in

accordance with the Vishay Micro-Measurements Application Note B-129-8 (Vishay 2009)

[23]. A static load was successively applied to the specimen. During the static test, specimens

were loaded within the elastic range and the monotonic load was gradually increased to its

maximum value in several equal-sized load increments/steps. Strain increments at the gage

locations were recorded for each load increment in order to determine the strain curve from

the strain gage reading. This curve indicates the material strain distribution for trapezoidal

ribs with time, as illustrated in Fig. 12.

The actual strain histories experienced by the model components were directly measured by

strain gages in the areas of concern where the stresses were likely to have important values. A

key component in bridge fatigue evaluation is the accurate determination of strain values; that

is, induced fatigue cracks. Compared to analytical methods, field testing provides the most

accurate approach, as no assumptions need to be made for load distribution uncertainties, such

as unintended composite action among structural components, contributions of nonstructural

members, various connection stiffness values, and steel rib behavior in tension [24].



In this case, a load of 100 kN was of interest in order to create a simulation with the results

obtained from FE method (ABAQUS) program, with the same load applied at each gage

10

location on the specimens. This proved to be an acceptable approach to the obtained

deformation results, as illustrated in Fig. 13.


4.2.2. Fatigue Data Measurement

Following completion of the static calibration load tests, all data were analyzed, and it was

decided to conduct dynamic calibration with test parameters. The experimental fatigue

loading formed a sinusoidal shape with only a compression part. Cyclic fatigue loads were

applied for all specimens, so that the minimum load was one tenth of the maximum (

R=Pmin ⁄ Pmax=0.1¿, and the maximum applied load was 100 kN. Cyclic loading was

applied using a sinusoidal function with a constant frequency of 2.0 Hz and load control. The

fatigue tests were terminated once the prescribed failure criteria of vertical rigidity were

reached. The specimens were tested by using these parameters until failure occurred. Crack

length readings with bare eyes were recorded approximately every 5000 cycles for each

specimen. The test was run continuously, as illustrated in Fig. 14. Therefore, the important

aspect of the fatigue analysis was to determine to what extent the crack grew under each cycle

of applied loading with this information, and counting the number of cycles that would lead to

failure could be carried out [25].


4.3. Fatigue Failure Observations

4.3.1. Failure Mode

Fatigue cracks at the RR butt weld connections were mainly caused by out-of-plane bending,

which is mode I loading. The failure modes of the five specimens with initial cracks were

11

similar, represented by an opening mode at the bottom of the ribs and indicated by local

displacements that are symmetric with respect to the x-y and x-z planes, and the two fracture

surfaces were displaced perpendicular to one another in opposite directions [26]. Fig. 15

illustrates the typical failure modes of the five specimens with initial cracks, where fatigue

cracks were all observed on the butt weld connections between the rib and connection plate,

commonly near the center of the rib arc.


The crack initiation and growth process can be summarized as the following three stages:

Stage I: An initial crack of 50 mm long and 6 mm deep was manually created at the center arc

of the butt weld on one side of the connection from the bottom of the rib for the five

trapezoidal specimens by using a cutting tool, after determining the exact length and depth

required, as shown in Fig. 16. The main reason for implementing the initial crack may be

considered as reducing the dynamic calibration test time. It was found that the control

specimen reached 1.2 million cycles, which is more than 10 consecutive days of continuous

examination without any noticeable cracks, following which the test was stopped. The

secondary reason is that it may be easier to investigate the propagation characteristics of

possible rib cracks. Therefore, the crack determination became more accurate and they could

be observed with the naked eye. Because the study focused on the areas around the initial

crack tips, these were manually adjusted according to the crack growth during the entire test

process, by using either the dye-penetrant or dynamic stiffness crack detection method, both

of which were implemented in this study. The crack initiation locations are consistent with the

highest stress concentration regions that were identified from the static calibration test and

computer simulation.

12


Stage II: As the cyclic loading continued, the crack steadily propagated away from its two

tips, up to the outer tip near the rib wall end, and finally reached the rib edge. The crack

growth paths always progressed along the butt weld toes on the rib. The crack growth rate

during this stage was slow compared to other states, and the inner tip propagated faster than

the outer tip, owing to the direct effect of cyclic loading. In this propagation stage, the cracks

became more visible to bare eyes by using dye-penetrant crack detection, as illustrated in Fig.

17, as their opening widths were gradually enlarged. The test was monitored and the specimen

was inspected for possible fatigue cracking at regular intervals. In the dye penetrant method,

the procedure recommendations were firstly achieved by spraying a BT68 spray type onto the

specimen on the parts where cracks were expected to occur, the function of which is to

penetrate inside the cracks rapidly. Repeated spraying using a BT69 spray type served to

clean only the surface part that was sprayed with BT68, following which it was wiped dry

with a cotton cloth. The use of these two sprays made the cracks visible to human eyes, and

the method was applied without stopping the test. This stage came to an end when cracks

began to appear on the top surface of the butt weld connections, and the maximal crack

opening displacements on the butt weld connection of the rib bottom surface were found to be

up to 1 mm.


Stage III: The outer tip of the crack had to propagate along the butt weld thickness, and finally

ran through it. The measured crack depth at this location increased rapidly as the cyclic

13

loading was repeated, while the crack inner tip continued to extend forward at a high speed.

The crack outer tip, which ran through the weld thickness, was extended in the vertical

direction perpendicular to the loading axis (towards the inner tip) step by step and increasing

crack parts ran through the weld. As a result, a visible longitudinal crack parallel to the weld

formed on the wall and bottom surface of the rib plate, and the measured opening widths were

up to 5 mm, as illustrated in Fig. 18.


Furthermore, the crack inner tip continued its propagation in the original direction (towards

the un-cracked edge of the rib plate). The crack growth speed during this stage was obviously

higher than the previous ones, as the crack tips were less constrained than ever. This final

stage ended when the specimen totally cracked and all the connections failed owing to tearing

of the rib plates.

4.3.2. Fatigue Crack Growth

The variations in crack length, as observed on the rib bottom surfaces and butt weld

connections during the test, are illustrated in Figs. 17 and 18. It is evident that the cracks

continued to increase in vertical length from initiation until the end of the test, and there were

no obvious distinguishing boundaries between the three previously described cracking stages.

The crack length growth rate (curve slopes in Fig. 19) also increased continuously during the

test, which means that the stress intensity factors at the crack tips increased as the crack length

extended [27]. Generally, during the early stage, approximate linear increases in crack length

were observed with the number of cycles before the crack outer tip arrived at the top rib edge.

Thereafter, the crack growth rate began to improve steadily as the cyclic loading continued,

and the crack extended rapidly in its vertical direction until the final length accumulated to

14

several times the edge crack length.


At a certain number of cycles (7500), the crack began to propagate at 2 mm long, and

cracking location investigations were performed using the dye-penetrant method, as

mentioned previously, following which the cracks were monitored and registered manually

versus every 5000 cycles until the specimen was exposed to fatigue issues. The crack growth

began from both surface edges of the initial crack, where the maximum bending stress was

generated and penetrated to the weld thickness center. Once the crack began to penetrate to

the weld thickness middle, the crack growth curve at the center of the weld connections was

almost parallel to the weld stretch shape, indicating that the crack growth rate at the center

and weld stretch shape were nearly identical. The measured total crack lengths and cycle

number fractures for each specimen are listed in Table 3.


4.4. Experimental Results and Discussion

4.4.1. Fatigue Life

As failure was predicted to occur at the weld connections, all specimens were tested at a

nominal stress range of 60.772 MPa, taken at the center of the butt weld connection from the

rib bottom side with the 100 kN applied loading. Where the stress range was too small, the

maximum stress state was taken from the actual state and not selected as higher. Each of the

specimens with initial cracks exhibited significantly shorter fatigue lives than the control

specimen tested within the same stress range. This effect is illustrated in Fig. 20, which

indicates the cumulative number of cycles for each specimen in the form of an S–N diagram

with the AASHTO (2007) fatigue design curves. The S–N curve is a relation between the

stress range under constant amplitude loading and the number of cycles until fracture [28].

15

Specimens with initial cracks underperformed the AASHTO Category E' fatigue curve by a

significant margin, which may be explained by the lowest fatigue strength quality of the

welds alone at the bottom of the ribs.


4.4.2. Crack Propagation Determination with Dynamic Stiffness Data

The instantaneous or dynamic stiffness of the specimens was a parameter inferred from direct

test measurements. Herein, the dynamic stiffness is defined as the change in the applied load

divided by the change in deflection of the specimen for each recorded fatigue cycle:

K dyn=∆ P ⁄ ∆ y (1)

where ∆ P is the change in applied load over one fatigue cycle and ∆ y is the change in

deflection over one fatigue cycle. The decrease in dynamic stiffness of the specimen served as

an indication of the change in the specimen response to loading. This change was a result of

the initiation and propagation of a crack in the model. As discussed in section 4.3.2, the

fatigue crack growth in a specimen was expected to cause a decrease in dynamic stiffness, as

recorded by the controller. The data used to calculate the dynamic stiffness were measured by

the Instron testing machine and recorded by the controller. All data were plotted using an

Excel spreadsheet. In order to display the subtle changes in dynamic stiffness as effectively as

possible, each plot indicates the narrowest dynamic stiffness range that the data recorded for

that specimen would allow. As illustrated in Fig. 21, the dynamic stiffness generally began to

decrease at a certain point following the initiation of crack propagation for specimens with

initial cracks. The decrease in dynamic stiffness of a specimen did not coincide perfectly with

the visual crack detection. In general, the dynamic stiffness did not begin to decrease until

approximately 7500 cycles after crack propagation, as detected using the dye penetrant.

Changes in dynamic stiffness recorded for the specimens that developed fatigue cracks

16

appeared to predict the fatigue cracking onset accurately. However, when testing specimens at

a nominal weld toe stress of 60.772 MPa, problems with this crack detection method became

apparent.

[Figures 21 about here]

A decrease in dynamic stiffness was demonstrated to be a direct indication of fatigue crack

growth. The fatigue crack propagation could be predicted by recording and monitoring

changes in the dynamic stiffness of a specimen throughout the fatigue test duration of a

fatigue test. Moreover, Fig. 22 presents the stiffness values versus every cycle during cycling

load until fatigue. It was observed throughout the cycling load that the change in the stiffness

value was small and almost a straight line, but the value suddenly changed downwards, which

means that the specimens with initial cracks failed [29]. However, the stiffness line for the

control specimen still does not exhibit any slope change, indicating that no noticeable cracks

appeared during the test, as illustrated in Fig. 22.

[Figures 22 about here]

5. Conclusions

The conclusions below were established by means of the FE model analysis and experimental

tests.

(1) The principal stresses on the bottom side of the trapezoidal rib, particularly at the

connection region, were higher than in other parts of the structural component; thus, fatigue

cracks appeared from the bottom of the ribs.

17

(2) The maximum deformations on the trapezoidal rib cross-section occurred closer to the

bottom side of the rib. This demonstrates that the rib wall exhibits out-of-plane bending

through the transverse directions. Following crack appearance, they easily propagated as a

result of out-of-plane bending.

(3) All specimens with initial cracks developed fatigue cracks at the center of the transverse

welds where the initial crack developed. In these specimens, the cracks propagated at either

the center or one end of the initial transverse weld toe and extended into the web in an

orthogonal direction to the longitudinal axis of the specimens. Once a crack had propagated in

these specimens, the crack growth rate increased exponentially, until yielding of the reduced

web section resulted in failure.

(4) During fatigue testing, it was concluded that the most effective method for detecting the

fatigue crack propagation onset was the modified dye-penetrant method. Another less

effective method that was employed to detect the fatigue crack propagation onset was the

dynamic stiffness method.

(5) By comparing the lives of fatigue specimens, it was concluded that specimens tested at the

center of the transverse weld connection from the bottom side of the trapezoidal rib provided

an effective means of predicting the fatigue life. The fatigue life of specimens with initial

cracks tested at a nominal weld toe stress range of 60.772 MPa achieved a level at or below

the AASHTO Category E' detail, while the control specimen exhibited behavior at or above

the curve expected for the AASHTO Category E detail.

(6) The numerical analysis and experimental test data exhibited very strong agreement.

Furthermore, although the stress range was low, the fatigue cracks propagated easily.

(7) One apparent difference between actual situations and the experimental results is that

fatigue cracks that propagated following the weld direction reached the deck plate, while in

18

actual situations fatigue cracks originating from the weld direction may be one third of the

weld height. The reason for this may be that in real situations, the rib is continues, while in

the experimental specimens it is only 1500 mm in length.

Acknowledgements

All of the tests were performed using the Instron machine, which was provided by project

number TR10/15/YNK/0034 from the İstanbul Development Agency under the project name

“Sönümleyici Mesnet Test Merkezi”.

References

[1] Liu R, Wang B, Liu Y. Distortion induced fatigue of deck plate at rib intersection with

diaphragm in orthotropic steel deck. International Journal of Steel Structures

2015;15:623–32. doi:10.1007/s13296-015-9009-3.

[2] Ju X, Tateishi K. Fatigue Crack Behavior at Rib-To-Deck Weld Bead in Orthotropic

Steel Deck. Advances in Structural Engineering 2014;17:1459–68. doi:10.1260/1369-

4332.17.10.1459.

[3] de Freitas ST, Kolstein H, Bijlaard F. Sandwich system for renovation of orthotropic

steel bridge decks. Journal of Sandwich Structures and Materials 2011;13:279–301.

doi:10.1177/1099636210386848.

[4] Sim H-B, Uang C-M, Sikorsky C. Effects of Fabrication Procedures on Fatigue

Resistance of Welded Joints in Steel Orthotropic Decks. Journal of Bridge Engineering

2009;14:366–73. doi:10.1061/(ASCE)1084-0702(2009)14:5(366).

[5] Sim H-B, Uang C-M. Stress Analyses and Parametric Study on Full-Scale Fatigue

Tests of Rib-to-Deck Welded Joints in Steel Orthotropic Decks. Journal of Bridge

19

Engineering 2012;17:765–73. doi:10.1061/(ASCE)BE.1943-5592.0000307.

[6] Xiao ZG, Yamada K, Ya S, Zhao XL. Stress analyses and fatigue evaluation of rib-to-

deck joints in steel orthotropic decks. International Journal of Fatigue 2008;30:1387–

97. doi:10.1016/j.ijfatigue.2007.10.008.

[7] Ya S, Yamada K, Ishikawa T. Fatigue Evaluation of Rib-to-Deck Welded Joints of

Orthotropic Steel Bridge Deck. Journal of Bridge Engineering 2011;16:492–9.

doi:10.1061/(ASCE)BE.1943-5592.0000181.

[8] Battista RC, Pfeil MS, Carvalho EML. Fatigue life estimates for a slender orthotropic

steel deck. Journal of Constructional Steel Research 2008;64:134–43.

doi:10.1016/j.jcsr.2007.03.002.

[9] Pfeil MS, Battista RC, Mergulhão AJR. Stress concentration in steel bridge orthotropic

decks. Journal of Constructional Steel Research 2005;61:1172–84.

doi:10.1016/j.jcsr.2005.02.006.

[10] Connor R. Influence of Cutout Geometry on Stresses at Welded Rib-to-Diaphragm

Connections in Steel Orthotropic Bridge Decks. Transportation Research Record:

Journal of the Transportation Research Board 2004;1892:78–87. doi:10.3141/1892-09.

[11] Connor RJ, Fisher JW. Consistent Approach to Calculating Stresses for Fatigue Design

of Welded Rib-to-Web Connections in Steel Orthotropic Bridge Decks. Journal of

Bridge Engineering 2006;11:517–25. doi:10.1061/(ASCE)1084-0702(2006)11:5(517).

[12] Steudle KT, Horsley J. AASHTO LRFD Bridge Design Specifications, 6th Edition -

June 2012 Errata 2012;20001.

[13] Choi J, Kim D. Stress Characteristics and Fatigue Crack Behaviour of the Longitudinal

20

Rib-to-Cross Beam Joints in an Orthotropic Steel Deck. Advances in Structural

Engineering 2008;11:189–98. doi:10.1260/136943308784466224.

[14] Oh CK, Hong KJ, Bae D, Do H, Han T. Analytical and experimental studies on optimal

details of orthotropic steel decks for long span bridges. International Journal of Steel

Structures 2011;11:227–34. doi:10.1007/s13296-011-2010-6.

[15] Xiao ZG, Yamada K, Inoue J, Yamaguchi K. Fatigue cracks in longitudinal ribs of

steel orthotropic deck. International Journal of Fatigue 2006;28:409–16.

doi:10.1016/j.ijfatigue.2005.07.017.

[16] Yokozeki K, Miki C. Fatigue evaluation for longitudinal-to-transverse rib connection of

orthotropic steel deck by using structural hot spot stress. Welding in the World 2016; 60:83–

92. doi:10.1007/s40194-015-0272-x.

[17] ASTM Int. Standard Test Methods for Tension Testing of Metallic Materials 1. Astm

2009:1–27. doi:10.1520/E0008.

[18] Brownjohn JMW, Dumanoglu AA, Severn RT. Ambient vibration survey of the fatih

sultan mehmet (second Bosporus) suspension bridge. Earthquake Engineering &

Structural Dynamics 1992;21:907–24. doi:10.1002/eqe.4290211005.

[19] ABAQUS Simulia 2009, Version 6.8-2. http://www.simulia.com.

[20] Wallerman H. Finite Element Design of Orthotropic Steel Bridge Decks 2015.

[21] Alemdar F. Evaluation of the Performance of Stress Measures for Longitudinal Ribs on

Orthotropic deck Using Finite Element Analysis 2015.

doi:10.13140/RG.2.1.2592.7769.

[22] American A, Standard N. Welding Code — 2004.

21

[23] Vishay. Surface preparation for strain gauge bonding: Application Note B-129-8

2009:1–7.

[24] Zhou YE. Assessment of Bridge Remaining Fatigue Life through Field Strain

Measurement. Journal of Bridge Engineering 2006;11:737–44.

doi:10.1061/(ASCE)1084-0702(2006)11:6(737).

[25] Liao J. Fatigue Damage in the Orthotropic Steel Deck with respect to the Trough‐to‐

Deck Plate Joint in between the Crossbeams 2011.

[26] Barsom J, Rolfe S. Fracture and Fatigue Control in Structures: Applications of Fracture

Mechanics, Third Edition. 1999. doi:10.1520/MNL41-3RD-EB.

[27] Yang M, Kainuma S, Jeong YS. Structural behavior of orthotropic steel decks with

artificial cracks in longitudinal ribs. Journal of Constructional Steel Research

2018;141:132–44. doi:10.1016/j.jcsr.2017.11.007.

[28] Ojamo T. Fatigue of the Orthotropic Layer Fiber Reinforced Composites 2012.

[29] Nijssen RPL. Fatigue life prediction and strength degradation of wind turbine rotor

blade composites. 2006.

List of Tables

Table 1 Tensile coupon test results

Table 2 Locations of strain gages in the model

Table 3 Total crack length versus number of cycles

22

List of Figures

Fig. 1. Fatigue crack in trapezoidal longitudinal rib:(a) initiate at the welded joint, and (b)

propagate to the rib wall.

Fig. 2. (a) Coupons while testing, and (b) coupon geometry following test.

Fig. 3. Trapezoidal rib details (dimensions in mm).

Fig. 4. (a) Finite element assembly mesh, and (b) mesh size at the target connections.

Fig. 5. Finite element modeling of trapezoidal rib connection plate.

Fig. 6. Maximum principal stress contours continuous trapezoidal rib side view

Fig. 7. Magnitude of displacement on trapezoidal longitudinal rib.

Fig. 8. Longitudinal paths applied in model: (a) path 1, (b) path 2, and (c) path 3.

Fig. 9. Experimental specimen details.

Fig. 10. Specimens during testing.

Fig. 11. Figure 12. Gage installation in field: (a) gage 1,2,3, and (b) gage 4,5.

Fig. 12. Strain gage distributions.

Fig. 13. Comparison of strain gage values of numerical and experimental approaches: (a) path 1,

(b) path 2, and (c) path 3.

Fig. 14. Run setup.

Fig. 15. Failure mode at bottom of rib.

23

Fig. 16. Initial crack.

Fig. 17. Detection of crack propagation by using a dye-penetrant method.

Fig. 18. Extended fatigue crack thickness along butt weld connection.

Fig. 19. Crack length versus number of cycles (R=0.1).

Fig. 20. S-N diagram illustrating results of six trapezoidal ribs testing in AASHTO (2007)

fatigue curves.

Fig. 21. Dynamic stiffness versus cycles for specimens with initial cracks.

Fig. 22. Dynamic stiffness versus cycles for control specimen.

Tables

Table 1

Tensile coupon test results

Table 2

Strain gage locations in the model

24

Coupon No E, MPa Fy, MPa Fu, MPa

RC-01 192679 268.28 392.53

RC-02 190652 269.72 390.97

RC-03 201607 294.34 397.34

Average: 194979 277.45 393.61

DC-01 199054 275.88 413.76

DC-02 197896 291.31 440.70

DC-03 202573 301.58 428.40

Average: 199841 289.59 427.62

Table 3

Total crack length versus number of cycles

FIGURES

25

Gage No Approximate Position

G1 5 [mm] from butt weld joints


G3 Center of model 50 [mm] from beginning of model edge

G4 Center of model 750 [mm] from beginning of model edge


Specimen ID No. of Cyclesto Failure

Total Crack Length [mm]

Specimen 1 284555 470

Specimen 2 217260 410

Specimen 3 63288 310

Specimen 4 209529 380

Specimen 5 143467 350

Control specimen Runout No crack

Fig. 1. Fatigue crack in trapezoidal longitudinal rib:(a) initiate at the welded joint, and (b) propagate to the rib wall.

Fig. 2. (a) Coupons while testing, and (b) Coupon geometry following test.

26

(a) (b)

(a) (b)

Fig. 3. Trapezoidal rib details (dimensions in mm).

Fig. 4. (a) Finite element assembly mesh, and (b) mesh size at the target connections.

Fig. 5. Finite element modeling of trapezoidal rib connection plate.

27

Mesh size: 5[mm]

Butt WeldConnection Plate

Deck Plate

Trough Rib

Plate under Load

FixtureHolder

Target connections

(a) (b)

Fig. 6. Maximum principal stress contours continuous trapezoidal rib side view

Fig. 7. Magnitude of displacement on trapezoidal longitudinal rib.

28

29

(a)

Path 1 G1G33

G2Path 2

(b)

Fig. 8. Longitudinal paths applied in model: (a) path 1, (b) path 2, and (c) path 3.G =Indicate the strain gage.

Fig. 9. Experimental specimen details.

30

G4Path 3 G5

(c)

Fig. 10. Specimens during testing.

Fig. 11. Gage installation in field: (a) gage 1,2,3, and (b) gage 4,5.

31

(a) (b)

G2G1G3 G4 G5

11:49:12 11:49:55 11:50:38 11:51:21 11:52:04 11:52:48 11:53:31-100

0

100

200

300

400

500

600

700G1G2G3G4

Time

Stra

in (m

m/m

m)

Fig. 12. Strain gage distributions.

0 200 400 600 800 1000120014001600

-0.00005

0

0.00005

0.0001

0.00015

0.0002

0.00025

0.0003

0.00035Path1G1"Analytical"G3"Analytical"G1"Experimental"G3"Experimental"

Distance (mm)

Stra

in (m

m/m

m)

020

040

060

080

010

0012

0014

0016

000

0.00005

0.0001

0.00015

0.0002

0.00025

0.0003

0.00035Path2G2"Analytical"G2"Experimental"

Distance (mm)

Stra

in (m

m/m

m)

32

(a) (b)

0 200 400 600 800 1000 1200 1400 1600-0.00005

0

0.00005

0.0001

0.00015

0.0002

0.00025

0.0003

0.00035Path3G4"Analytical"G5"Analytical"G4"Experimental"G5"Experimental"

Distance (mm)

Stra

in(m

m/m

m)

Fig. 13. Comparison of strain gage values of numerical and experimental approaches: (a) path 1, (b) path 2, and (c) path 3.

Fig. 14. Run setup.

33

(c)

Fig. 15. Failure mode at bottom of rib.

Fig. 16. Initial crack.

34

Opening Mode

50 [mm] initial crack

Fig. 17. Detection of crack propagation by using a dye-penetrant method.

Fig. 18. Extended fatigue crack thickness along butt weld connection.

35

0 50000 100000 150000 200000 2500000

50

100

150

200

250

300

350

400

450

specimen 1Specimen 2Specimen 3Specimen 4Specimen 5

Number of cycles

Cra

ck le

nght

(mm

)

Fig. 19. Crack length versus number of cycles (R=0.1).

10000 100000 1000000 10000000 10000000010

100

1000Specimen 1

Specimen 2

Specimen 3

Specimen 4

Specimen 5

Control specimen

Number of cycles ,N

Stre

ss r

ange

(MPa

)

ABB'

C

C'

D

EE'

B'

&

C'&

Fig. 20. S-N diagram illustrating results of six trapezoidal ribs testing in AASHTO (2007) fatigue curves.

36

= Indicate run-out

0 50000 100000 150000 200000 2500000

10

20

30

40

50

60

Specimen 1Specimen 2Specimen 3Specimen 4Specimen 5

Number of Cycles

Dyn

amic

stiff

ness

(N/m

m)

Fig. 21. Dynamic stiffness versus cycles for specimens with initial cracks.

0 200000 400000 600000 800000 1000000 1200000 14000000

10

20

30

40

50

60

Control specimen

Number of Cycles

Dyn

amic

stiff

ness

(N/m

m)

Fig. 22. Dynamic stiffness versus cycles for control specimen.

37

files.transtutors.com€¦ · web viewfatigue crack behavior at rib-to-rib butt weld connection in...

Documents