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EFFECTS OF MULTIPLE DAMAGE-HEAT STRAIGHTENING REPAIRS ON STEEL BEAMS

Keith Kowalkowski

Staff Engineer

Ruby and Associates, P.C.

E-Mail: [email protected]

Amit H. Varma

Assistant Professor

School of Civil Engineering

Purdue University

E-Mail: [email protected]

Purdue University

School of Civil Engineering

550 Stadium Mall Drive

West Lafayette, IN 47907

Phone: (765) 496-3419

Dept. Fax: (765) 496-1105

Submission Date: January 19, 2007

Word count: Text + Tables + Figures = 5000 + 1000 + 1500 = 7500

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ABSTRACT

Experimental investigations were conducted to evaluate the effects of multiple cycles of damage followed by heat

straightening repair on the structural properties and fracture toughness of steel bridge beams. Six W24x76 (or

equivalent) beam specimens made from A7, A588, or A36 steel (two beam specimens of each steel type) were

subjected to three cycles of damage followed by heat straightening repair. In each cycle, the beam specimens were

damaged statically in weak axis bending, and repaired using several half-depth Vee heats and an external restraining

force. Material specimens were machined from the flanges of the damaged-repaired beam specimens and tested

according to ASTM standards. The material test results indicated that multiple damage-repair cycles do not have a

significant influence on the elastic modulus, ultimate stress, and surface hardness of the steels. Multiple damage-

repair cycles generally increase the yield stress and reduce the ductility of the steels. The charpy V-notch (CVN)

fracture toughness of A7 and A36 steel beams repaired with maximum heating temperature (Tmax) equal to 650°C

was lower than the fracture toughness of the corresponding undamaged steels. The CVN fracture toughness of A588

steel beams repaired with Tmax equal to 650°C, and the A36 steel beam repaired with Tmax equal to 760°C was much

higher than the fracture toughness of the corresponding undamaged steels. Material specimens that were removed

from locations closer to the flange-web interface consistently had lower fracture toughness and ductility.

Keywords: heat treatment, steel bridges, notch toughness, stress, strain, hardness, rehabilitation.

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INTRODUCTION

Over-height trucks occasionally (impact) collide with steel girders of highway bridges, causing significant structural

damage. Composite steel girder bridges subjected to this type of collision damage in Michigan are most frequently

made from A7 or A373 steel (1). The structural damage caused by the collision consists of permanent plastic

deformations and can be described using the damage characterization scheme developed by Avent and Mukai (2, 3),

which includes categories S (strong axis), W (weak axis), T (torsional), and L (local buckling) damage for steel

shapes. The composite steel girders typically undergo categories T and L structural damage due to collision. This

includes out-of-plane deformation of the bottom flange, and local buckling or denting at the point of impact.

Several state highway agencies use heat straightening to repair damaged steel girders. Heat straightening is

a cost-effective technique for repairing structural damage caused by collisions. It consists of applying heating in

specific patterns (Vee, strip, or spot heats) and locations to reverse the plastic deformations and repair the damage.

Heating is applied using oxygen-fuel torches, while maintaining the steel surface temperature below 650°C (1200°F)

for mild steels and 595°C (1100°F) for quenched and tempered steels (2, 3). Combinations of Vee, strip, and spot

heats are used to repair the damaged composite steel girders, where Vee heats are used to repair the out-of-plane

deformations of the flange, strip heats are used to repair the web, and spot heats are used to repair the local buckles

and dents (2, 3, 4). Several heating cycles are applied until the residual deformations reach within tolerance levels

and the damage is repaired.

Significant research has been conducted on the heat-straightening repair of damaged steel shapes (4, 5, 6),

and the effects of damage and repair on steel beams (6, 7, 8). Based on this research, heat-straightening repair

guidelines have been developed by Shanafelt and Horn (9) and Avent and Mukai (2). These guidelines include:

techniques for measuring damage, recommendations for heat straightening patterns, limits on restraining forces

during repair, maximum heating surface temperatures, and tolerances for residual deformations after repair. These

guidelines are used by some state departments of transportations (DOTs) across the U.S., while other DOTs use in-

house research, and some do not have any heat repair guidelines (1).

In some cases, the same steel bridge girder is subjected to multiple damage and heat straightening repair

events over its life. This raises concerns regarding their condition and serviceability after multiple damage-repair

events. Current guidelines indicate that the same steel girder should not be subjected to more than two damage-heat

straightening repairs. This recommendation is based on research conduced by Avent et al. (7, 10) Currently, DOTs

have significant interest in developing rational guidelines for evaluating the acceptability of steel girders subjected

to multiple damage-repair events (1). This paper presents the results of a research project focusing on this topic.

BACKGROUND

Significant research has been conducted on the heat-straightening repair of steel girder bridges. Researchers have

conducted experimental investigations to determine the deformations and residual stresses produced by heat

straightening (4, 5, 6, 8). The results from these studies were used to develop empirical equations and guidelines for

conducting heat straightening in the field (2). The effects of heat straightening on the structural properties (yield and

ultimate stress, ductility, and hardness) of bridge steels have also been studied by several researchers (7, 11, 12). The

results from these studies indicate that heat straightening increases the yield and ultimate stress, and decreases the

ductility of repaired steels.

Avent et al. (7) also investigated the effects of multiple damage-heat straightening repair cycles on the

structural properties of A36 steel W6x9 beams. They subjected the steel beam specimens to 1, 2, 4, and 8 damage-

repair cycles. The beams were damaged by weak axis bending (Category W), and repaired using restraining

moments equal to 50% of the plastic moment (0.50Mp) and Vee heats in the damaged regions (flanges). Tension

coupons were taken from damaged-repaired regions near the apex, center, and open end of the Vee, and tested

according to ASTM standards. The results indicated that the damage-repair cycles increase the yield and ultimate

stress, and decrease the ductility of the steel. The reduction in ductility was moderate after 1-2 cycles, but significant

after 4-8 damage-repair cycles. Hence, Avent et al. (7) suggested a limit of two damage-repair cycles for steel

girders. Avent and Fadous (10) subjected a composite W24x76 A36 steel beam to four damage-repair cycles. During

the fourth cycle, a crack formed on the convex side of the web yield line. This further reinforced the

recommendation to limit steel bridge girders to two damage-repair cycles.

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RESEARCH PROJECT

The authors (13) have conducted a research project focusing on the effects of multiple damage-heat straightening

repair cycles on the structural properties (elastic modulus, yield and ultimate stress, ductility) and fracture toughness

of common (ASTM A7, A36, and A588) bridge steels. This project consisted of two phases.

The first phase conducted laboratory-scale tests on steel plate specimens by subjecting them to multiple

damage-heating repair cycles. The damage and repair parameters included were: (i) the inelastic damage strain

(!d=30, 60, or 90 times the yield strain), (ii) the restraining stress ("r=25 or 50% of the yield stress), and (iii) the

number of damage-repair cycles (Nr=1, 2, 3, 4, or 5 cycles). The effects of each parameter were evaluated by

varying it (within its range), while holding all other parameters constant. Most of the A7, A588, and A36 steel

specimens were repaired with the maximum heating temperature (Tmax) limited to 650°C (1200°F). Some A36 steel

specimens were repaired with Tmax equal to 760°C (1400°F) or 870°C (1600°F). A total of ninety steel plate

specimens were tested. The experimental results and major findings from the phase-1 investigations are presented in

detail elsewhere (13, 14).

The second phase conducted large-scale experimental investigations on steel beam specimens to further

evaluate the effects of multiple damage-heat straightening repair cycles on the structural properties and fracture

toughness of steel beams, and to evaluate and verify the findings from phase-1. This paper presents the experimental

investigations and findings from phase-2 of the project.

EXPERIMENTAL INVESTIGATIONS – PHASE 2

Test Matrix

Table 1 shows the test matrix for the experimental investigations of phase-2. It includes the nomenclature for the

beam specimens along with the details of damage and repair parameters. The nomenclature consists of the steel type

followed by the specimen number, for example, A7-Beam-1, A36-Beam-2, etc. In Table 1, !d is the maximum

damage strain, !y is the undamaged steel yield strain, #p is the plastic displacement after damage, Mr is the maximum

restraining moment, Mp-y is the weak-axis plastic moment capacity of the beam, Tmax is the maximum temperature

achieved during heating, and Nr is the number of multiple damage-repair cycles.

As shown in the Table, two specimens of each steel type (A7, A588, and A36 steel) were tested. All the

specimens were subjected to three damage-repair cycles because that was the limit recommended by phase-1 (13,

14) for A7 and A36 steel, and needed further evaluation and verification in phase-2. Typically, one major damage-

repair parameter was varied between the two specimens (of each steel type), while the remaining parameters were

held relatively constant. The parameter that was varied between the two beam specimens of each steel type was

based on the phase-1 findings presented in (13, 14):

• The damage strain was the major parameter varied between the two A7 steel beam specimens. A7-Beam-1 was

subjected to the three cycles of damage-repair with !d equal to 30 !!y. A7-Beam-2 was subjected to three cycles

of damage-repair with !d equal to 90, 60, and 60!y. A7-Beam-1 was repaired with Mr equal to 0.25, 0.50,

0.25Mp-y for cycles 1, 2, and 3, respectively. A7-Beam-2 had to be repaired with Mr equal to 0.50Mp-y because

numerous Vee heats were required to repair the larger damage strains. The maximum heating temperature (Tmax)

was equal to 650oC (1200°F) for all three cycles.

• The restraining moment was the parameter varied between the two A588 steel beam specimens. Both specimens

were subjected to three cycles of damage-repair with !d equal to 40, 20, and 20!!y for cycles 1, 2, and 3,

respectively. Tmax was equal to 650°C (1200°F) for all three cycles. A588-Beam-1 was repaired with Mr equal to

0.25Mp-y, and A588-Beam-2 was repaired with Mr equal to 0.50Mp-y for all three damage-repair cycles.

• The heating temperature was the parameter varied between the two A36 steel beam specimens. Both specimens

were subjected to three cycles of damage-repair with !d equal to 30!y, and Mr equal to 0.25, 0.50, and 0.25 Mp-y

for cycles 1, 2, and 3, respectively. A36-Beam 1 repaired with Tmax equal to 650oC (1200°F), and A36-Beam 2

was repaired with Tmax equal to 760oC (1400°F). The smaller damage strain (!d=30!y) was used because it was

more detrimental than larger damage strains according to the phase-1 findings.

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Testing Approach

The authors visited three heat straightening repair sites to determine the heat straightening techniques, limitations,

and procedures employed by the Michigan Department of Transportation (MDOT) Statewide Bridge Crew (SBC).

In all three cases, the fascia girders of simply supported composite steel bridges were damaged by collision with

over-height trucks and repaired by heat straightening. The composite steel girders exhibited large out-of-plane

deformations of the bottom flange. Figure 1(a-b) shows the damage for two of the three bridges.

The heat-straightening repair procedure was also similar for all three cases. It consisted of applying an

external restraining force in the direction opposite to the impact force using the equipment shown in Figure 2(a).

Strip heats were applied to the top of the web just below the concrete slab as shown in Figure 2(b). Several half-

depth Vee heats were applied to the bottom flange using oxy-acetylene torches as shown in Figure 2(c). The half-

depth Vee heats were spaced along the damaged (plastic) regions as shown in Figure 2(d). These Vee heats had a

constant base width of 125 mm, resulting in Vee angles of 45-55°. The authors documented each repair performed

by the SBC and observed some overheating in the range of 690-760ºC (1300-1400°F) in some Vee heats. This repair

procedure is used most commonly by the SBC and is similar to the one used by Avent and Mukai (2, 3).

The testing approach was designed to approximate the damage and heat straightening repair procedures

used by the SBC. The out-of-plane flange deformations of a composite steel girder (see Figure 1) are similar to the

flange deformations of a steel beam damaged in (category W) weak axis bending (2, 3). The heat-straightening

repair of beam flanges damaged in weak axis bending consists of several partial depth Vee heats spread along the

length of flanges, which is also similar to the heat-straightening repair of the damaged flanges used by the SBC (2,

3). Based on these similarities, the steel beam specimens were subjected to category-W static damage and repair for

the experimental investigations.

This testing approach facilitates much better control of the damage strain and the restraining stress in the

beam specimens during damage-repair cycles. Additionally, in weak axis bending, both flanges of the steel beam are

subjected to identical damage and repair histories, which provides twice the material for obtaining samples for

material testing. Material specimens for tension and fracture toughness testing can be obtained from both flanges,

thus providing enough data for confidence in the results. The major limitation of the testing approach is that it does

not simulate the impact damage of composite steel girders, which is difficult to quantify in the field and duplicate in

the laboratory with the facilities available.

Test Setup

Figure 3 shows photographs of the test setup used to subject the beam specimens to damage (weak axis bending)

followed by heat straightening repair. The beam specimens were roller supported at the ends, and subjected to

concentrated forces at the mid-span. Figures 3(a) and (b) show photographs of a beam specimen before and after

damage, respectively. The various components of the test setup are identified in these Figures. Additional details of

the test setups are presented elsewhere in (13, 15).

Figure 4 shows a schematic of the loading frame used to apply concentrated forces at the beam midspan in

both the upward and downward directions. The concentrated forces were applied using a double-acting hydraulic

actuator that was bolted and post-tensioned to the laboratory strong floor. The actuator applied the damaging

(upward) force by pushing the loading frame against the flanges of the beam specimen. The restraining force was

applied as two equal concentrated forces acting at 120 mm. on either side of the beam midspan. Additional details of

the loading frames are presented in (13, 15).

Test Specimens

Six beam specimens were fabricated from A7, A588, and A36 steels (two of each steel type). The A7 beam

specimens were fabricated from a decommissioned W24x76 A7 steel beam (7.3m long) provided by MDOT. A

portion was removed from the beam mid-length to fabricate material coupons for evaluating the undamaged A7 steel

properties, and two 2.9m long A7 beams specimens were fabricated from either side of this central portion.

The A588 and A36 beam specimens were built-up sections with details and dimensions similar to the

W24x76 A7 steel beam specimens. The flange plates were 230mm wide and 19mm thick, and the web plates were

530mm wide and 13mm thick. The flange plates were attached to the web plates using 10mm. fillet welds. Portions

of the flange and web plates were retained for fabricating material coupons to determine the undamaged A588 and

A36 steel material properties.

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Instrumentation

Several sensors were used to monitor the experimental behavior during damage and repair. Two pressure

transducers were used to measure the actuator pressures and calculate the applied forces. The midspan deflections

were measured using two displacement transducers that were placed on the opposite flange plates of the beam

specimens. The quarter point deflections were also measured using two displacement transducers that were placed

on either side of the beam midspan (halfway to supports). The end rotations of the beam specimens were measured

using two rotation meters. The time-temperature (T-t) response of the heated steel was measured using a non-contact

infrared thermometer. Figure 3(a) includes some details of the instrumentation layout for the beam specimens.

Three longitudinal strain gages were also bonded to the flanges of the beam specimen at the mid-span.

These strain gages were used to monitor the maximum damage strain (!d) during the first damage cycle. They could

not be used in subsequent damage cycles because the heat straightening repair process produced small out-of-plane

distortions in the beam flanges. The mid-span deflections measured at the target !d in the first damage cycle were

used as the target displacements for the subsequent multiple damage cycles.

Heat Straightening Procedure

The Vee heat locations and distributions along the beam length were determined based on the plastic strain

distributions produced by the damage process. The inelastic force-displacement behavior of the beam and the

corresponding plastic strain distributions were calculated using simple numerical models. The beam length was

discretized into several stations, and the bending moment (M) at each station was computed using static force

equilibrium. The curvature ($) corresponding to the moment (M) at each station was obtained from the section M-$

response, which was computed using a fiber model of the cross-section and the measured undamaged steel uniaxial

stress-strain curves for the fibers. The station curvatures ($) were used to compute the plastic strain distributions

over the section and along the beam length.

Figure 5 shows an example of the plastic strain distribution calculated for specimen A7-Beam-2 subjected

to the damage strain of 90!y in the first cycle. It shows the plastic strain distribution as a function of the distance

from the midspan. Figure 6 shows the location, distribution, and nomenclature of multiple overlapping Vee heats

that were used to repair the damaged specimens. The depth and width of each Vee heat were approximately equal to

115 mm, which corresponds to a Vee angle of 53°. Figure 5 superimposes the Vee heat locations on the plastic strain

distributions to illustrate that the Vee heats were located and distributed along the beam length to repair most of the

material subjected to plastic strains (damage).

The heat straightening repair was conducted by subjecting each beam specimen to several cycles of

restraining force (producing moment Mr) and heating. Each heating cycle consisted of applying two simultaneous

Vee heats directly across from each other on opposite flanges using two oxy-acetylene torches. In each Vee, a spot

heat was applied initially at the apex of the Vee until the target temperature (Tmax) was reached. The torch was then

moved slowly in a serpentine path from the apex to the open end of the Vee. The surface temperature of the steel

was monitored using a hand-held temperature-indicating device.

Several heating cycles were required to repair each damaged specimen. The Vee heats were alternated

between the various Vee locations (C, L1, R1, etc.) identified in Figure 6. More Vee heats were applied at the

locations with larger plastic strains, i.e., C, L1, L2, R1, and R2. The displacements and end rotations were used to

establish the completion of heat straightening repair. Vee heats were applied until the midspan displacement of the

beam specimen was within 1.5 mm (tolerance) of the displacement before damage. Figure 5 shows an example of

the number of Vee heats required at the various locations to complete the repair of specimen A7-Beam-2.

EXPERIMENTAL BEHAVIOR

Damage and Repair Behavior

Figure 7 shows the measured force-displacement behavior for all three damage cycles of specimen A36-Beam-1.

The behavior shown in Figure 7 is typical and representative of the damage behavior of all beam specimens. This

Figure shows that after each damage-repair cycle, the flexural stiffness of the beam decreases and it resistance

(strength) increases for larger displacements. This behavior is influenced by the residual stresses produced by the

damage and heat straightening repair processes (3, 4). The portions of the flanges subjected to high residual stresses

undergo yielding at lower loads, which results in reduced flexural stiffness. The portions of the flanges with

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significant residual strains undergo strain hardening immediately after yielding due to the prior damage-repair

excursions and history of the material.

Figure 8 shows the experimental behavior of specimen A7-Beam-1 during the third damage-heat straightening

repair cycle. It includes the measured surface temperatures, restraining forces, and the quarter and midspan

displacements as functions of time. The behavior shown in Figure 8 is typical and representative of the behavior of

all beam specimens during heat straightening repair. For clarity of presentation, Figure 8 focuses on the

experimental behavior measured during the first four Vee heats, which were applied sequentially to the Vee

locations C, R1, L1 and R2. The surface temperature-time curves indicate that the steel surface heats rapidly (in less

than one minute) to the target value (Tmax) and cools slowly (over 15 minutes) to room temperature before the next

Vee heat is applied. The applied restraining force increases slightly as the steel undergoes thermal expansion due to

heating. The applied restraining force decreases significantly as the steel undergoes yielding and then contraction

due to cooling. The displacement-time curves indicate that approximately 4-5 mm midspan deflection (repair) was

achieved during each Vee heat with restraining moment (Mr) equal to 0.50Mp-y.

Material Testing Procedures

Standard (ASTM) tests were conducted on material specimens fabricated from the flanges of the beam specimens

subjected to multiple damage-repair cycles. These included: (a) uniaxial tension tests (b) charpy V-notch (CVN)

fracture toughness tests, and (c) Rockwell hardness tests conducted according to ASTM standards (16, 17, 18). All

the CVN fracture toughness tests were conducted after cooling the CVN specimens to 4.4°C (40°F), which is the

temperature recommended for Zone 2 fracture toughness tests in (19).

The uniaxial tension coupons and the CVN specimens were fabricated from the two flanges of the

damaged-repaired beam specimens, which were identified as flanges A and B, respectively. Three tension coupons

were fabricated from the central Vee heated region within 65mm of the edge of flange A. Figure 10(a) shows the

original location of these tension coupons in flange A. It also identifies the nomenclature used to identify the tension

coupons. As shown in Figure 9(a), X, Y, and Z represent the outermost, center, and innermost coupon closest the

flange-web junction, respectively. These tension coupons had a gage length of 50mm, and their remaining

dimensions (width, thickness, etc.) were in accordance with (16).

Twelve CVN specimens were fabricated from the central region of flange B. Figure 9(b) shows the original

locations of these CVN specimens in the flange. As shown, the twelve CVN specimens were removed in three

columns of four specimens each. Each column of specimens was centered at one of the central Vee heats. The

nomenclature for identifying the CVN specimens consisted of the Vee heat location (C, L1, or R1) followed by the

numbers 1 to 4, where 1 is closest to the flange edge and 4 is closest to the flange-web interface. As shown in Figure

9(b), the CVN specimens were located within 65 mm of the flange edge (i.e., at least 50 mm from the flange web

interface) and fabricated from the mid-thickness of the flange plates. The notches were located along the centerline

of the Vee heats, and oriented to measure the through thickness fracture toughness. The Rockwell hardness tests

were conducted on the surfaces of CVN specimens C-2.

Standard (ASTM) tests were also conducted on material samples obtained from the flange plates (A and B) of

undamaged A7, A588, and A36 steel beams. The number and locations of the uniaxial tension coupons and CVN

specimens were similar to those shown in Figure 9. The results from the uniaxial tension tests were used to compute

the average elastic modulus (Eo), yield stress ("yo), ultimate stress ("uo), and ductility (% elongation eo) for the

undamaged steels. The results from the CVN impact tests at 4.4°C (40°F) were used to compute the mean fracture

toughness (FTo) and the 95% confidence interval of the fracture toughness for the undamaged steels. Rockwell

hardness tests were conducted on the surfaces of CVN specimen C-2 to compute the average surface hardness

(HRBo). These results for the undamaged A7, A588, and A36 steels are summarized in Table 2

Material Test Results

The uniaxial tension test results for coupons (X, Y, and Z) from flange A of the damaged-repaired beam

specimens were used to calculate the elastic modulus (E), yield stress ("y), ultimate stress ("u), and ductility (%

elongation, e) for the damaged-repaired steels. These results were normalized with respect to the corresponding

undamaged steel material properties (Eo, "yo, "uo, and eo), and the normalized values of E/Eo, !y/"yo, "u/"uo, and e/eo

are summarized in Table 3 for the six beam specimens. The normalized surface hardness values (HRB/HRBo) for the

damaged-repaired beam specimens are also included in Table 3.

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The fracture toughness test results for the CVN specimens from flange B of the damaged-repaired A7, A588,

and A36 steel beam specimens are presented in Tables 4, 5, and 6, respectively. Each table includes the fracture

toughness test results for the twelve CVN specimens (L1-1 to L1-4, C-1 to C-4, and R1-1 to R1-4) from both

specimens of the same steel type. The tables also include the average (Avg-H) of the fracture toughness values for

three CVN specimens at the same horizontal level, and the average (Avg-V) of the fracture toughness values for four

CVN specimens in the same vertical column. The fracture toughness values less than the AASHTO recommended

limit of 20 J at 4.4°C (40°F) are highlighted in bold. Additionally, each table also presents the normalized fracture

toughness values for the CVN specimens of both beam specimens. The normalized values are obtained by dividing

the measured fracture toughness values by the mean fracture toughness (FTo) of the undamaged steel from Table 2.

These normalized fracture toughness values eliminate any bias produced by the value (high or low) of the

undamaged steel fracture toughness, and indicate the effects of damage-repair cycles more clearly.

EXPERIMENTAL RESULTS

Evaluation of Material Test Results for A7 Steel

The normalized material properties in Table 3 indicate that the damage-repair cycles increase the yield stress of A7

steel to approximately 107-125% of the undamaged steel yield stress. The increase in yield stress was greater for

specimen A7-Beam-1, which was subjected to lower damage strains. For both beam specimens, the increase in the

yield stress was higher for the tension coupons closer to the flange-web interface. The damage-repair cycles do not

have a significant influence on the elastic modulus, ultimate stress, and surface hardness of the A7 steel specimens.

These values are within ± 10% of the corresponding undamaged A7 steel material properties.

The damage-repair cycles reduce the ductility of the A7 steel specimens to about 80-96% of the undamaged

steel ductility. This reduction is slightly greater for specimen A7-Beam-2, which was subjected to larger damage

strains. The results in Table 4 indicate that the fracture toughness values for specimen A7-Beam-1 were quite low.

Several fracture toughness values were less than the AASHTO recommended limit of 20 J. The toughness values

were lowest at the center (C) Vee heat, and they reduced further for CVN specimens closer to the flange-web

interface. The overall average fracture toughness of A7-Beam-1 was about 20% of the undamaged steel fracture

toughness. Most of the fracture toughness values (11 of 12) were less than 25% of the undamaged steel toughness.

The fracture toughness values for specimen A7-Beam-2 were reasonable. Only a few (3 out of 12) values

were less than the AASHTO limit of 20 J. The toughness values were lowest at the center (C) Vee heat and reduced

further for CVN specimens closer to the flange-web interface. The overall average fracture toughness of A7-Beam-2

was 105% of the undamaged steel fracture toughness. The CVN specimens close to the flange edge had very high

toughness values (approximately 107-284% of the undamaged steel toughness), and CVN specimens close to the

flange-web interface had very low toughness values (approximately 12-22% of the undamaged toughness).

These results indicate that damage strain is an important parameter for A7 steel specimens. Specimen A7-

Beam-1, which was subjected to smaller damage strains, had much lower fracture toughness values than specimen

A7-Beam-2. These findings are consistent with those from phase-1 of the study (13, 14).


The normalized material properties in Table 3 indicate that damage-repair cycles do not have a significant influence

on the yield stress, elastic modulus, ultimate stress, and surface hardness of A588 steel specimens. These values are

within ± 10% of the corresponding undamaged A588 steel material properties for both beam specimens. The yield

and ultimate stresses were slightly higher for the coupons (Z) closer to the flange-web interface. The damage-repair

cycles reduce the ductility of A588 steel specimens to about 79-103% of the undamaged steel ductility. The ductility

values were slightly higher for the coupons (X) closer to the flange edge, and they were consistently lower for the

coupons (Z) closer to the flange-web interface. The change in ductility was similar for both A588 steel specimens.

The results in Table 5 indicate that the fracture toughness values were very high for both A588 steel

specimens. The toughness values were always greater than the AASHTO requirement of 20 J, and in most cases

they were even greater than the undamaged steel fracture toughness (129.75 J). The toughness values were highest

for the CVN specimens closest to the flange edge. They decreased consistently with distance from the flange edge,

and were lowest for the CVN specimen closest to the flange-web interface. The overall average fracture toughness

of A588-Beam-1 and A588-Beam-2 were equal to 258% and 204% of the undamaged steel toughness, respectively.

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These results indicate that specimen A588-Beam-2, which was repaired using larger restraining moments

Mr, had slightly lower fracture toughness values than specimen A588-Beam-1. The restraining moment is a relevant

but not significant parameter for A588 steels. These findings are consistent with those from phase-1 (13, 14).


The results in Table 3 indicate that the damage-repair cycles increase the yield stresses and reduce the ductility of

the A36 beam specimens. The yield stresses of specimens A36-Beam-1 and A36-Beam-2 increased to about 107%

and 110-121% of the undamaged steel yield stress, respectively. The ductility of specimens A36-Beam-1 and A36-

Beam-2 reduced to about 88-77% and 79-89% of the undamaged steel ductility, respectively. The increase in yield

stress and reduction in ductility were more significant for specimen A36-Beam-2, which was repaired with higher

Tmax. The damage repair cycles do not have a significant influence on the elastic modulus, ultimate stress, and

surface hardness of the A36 beam specimens. These values are within ± 10% of the corresponding undamaged A36

steel material properties for both beam specimens.

The results in Table 6 indicate that the fracture toughness values for A36-Beam-1 were highly variable.

Most of the toughness values were greater than the AASHTO limit of 20 J, but some values were less than or close

to the AASHTO limit. The toughness values were typically higher for the CVN specimens closer to the flange edge,

and they decreased consistently with the distance from the flange edge. The average fracture toughness of A36-

Beam-1 was about 84%, and the lowest toughness values were about 22-33% of the undamaged steel toughness.

The fracture toughness values for A36-Beam-2 were much greater than the AASHTO recommended limit

of 20 J, but they were also quite variable. Most of the fracture toughness values (with one exception) were greater

than the undamaged steel fracture toughness (65.35 J). The fracture toughness values were higher for CVN

specimens closer to the flange edge, and they decreased consistently with the distance from the flange edge. The

average fracture toughness for specimen A36-Beam-2 was about 227% of the undamaged steel fracture toughness.

These results indicate that Tmax has a significant influence on the fracture toughness of A36 steel beam

specimens. Specimen A36-Beam-2 had much higher fracture toughness values than specimen A36-Beam-1. These

findings are consistent with those from phase-1 of the study (13, 14).

CONCLUSIONS

The results from the experimental investigations and the material tests indicate that multiple (three) damage-heat

straightening repair cycles do not have a significant influence on the elastic modulus (E), ultimate stress ("u), or

surface hardness (HRB) of A7, A588, or A36 steels, i.e., the values after three damage-repair cycles are within

±10% of the undamaged values.

The three damage-repair cycles increase the yield stress and reduce the ductility of A7 steel beam

specimens. The increase in yield stress is greater for specimens subjected to smaller damage strains. The reduction

in ductility is reasonable but not significant. The three damage-repair cycles reduce the fracture toughness of the A7

steel beam specimen subjected to smaller damage strains significantly, with several values less than the AASHTO

recommended limit of 20J at 4.4°C (40°F). The A7 steel beam specimen subjected to larger damage strains had

reasonable fracture toughness values. These findings are consistent with those from phase-1 of the study (13, 14).

The three damage-repair cycles do not have a significant influence on the yield stress or ductility of the

A588 steel beam specimens. The reduction in ductility is reasonable but not significant. After three damage-repair

cycles, the fracture toughness of the A588 steel beam specimens was much greater than the toughness of the

undamaged A588 steel. The A588 beam specimen repaired with the larger Mr had slightly lower fracture toughness

than the specimens repaired with smaller Mr. These findings are also consistent with those from phase-1 (13, 14).

The three damage-repair cycles increase the yield stress and reduce the ductility of the A36 steel beam

specimens. The increase in yield stress and reduction in ductility are more significant for the A36 beam specimen

repaired with the higher Tmax (760°C or 1400°F). After three damage-repair cycles, the fracture toughness values for

the A36 beam specimen repaired with lower Tmax (650°C or 1200°F) are quite variable, and in some cases less than

the AASHTO recommended limit of 20J. The fracture toughness values for the A36 beam specimen repaired with

higher Tmax (760°C) are also variable, but almost always greater than the fracture toughness of the undamaged A36

steel. These findings are also consistent with those from phase-1 of the study (13, 14).

10

LIMITATIONS

The major limitation of this study is that the testing approach does not account for the very localized effects of

impact damage, which are difficult to quantify in the field and implement in the laboratory. The damage

implemented in this research focused on weak-axis bending, and its similarity to the out-of-plane deformations of

composite beam flanges in real bridge structures. The research findings must be considered carefully before major

decisions or recommendations. Additional research is recommended to include the effects of realistic impact

damage, local denting and buckling, and to further investigate the effects of overheating above 650°C.

ACKNOWLEDGMENTS

This research was sponsored by the Michigan Department of Transportation (Roger Till–cognizant official) and

conducted using the laboratory facilities of Michigan State University. The data compilation and report preparation

were supported by Purdue University.

REFERENCES

1. Varma A, H., Kowalkowski, K.J., and Shingledecker, J. (2004). “Multiple Heat Straightening Repair of

Damaged Steel Bridges,” Proc. of the Transportation Research Board Meetings, TRB, NAS, [computer file].

2. Avent, R.R., and Mukai, D.J. (1998). Heat-Straightening Repairs of Damaged Steel Bridges - A Technical

Guide and Manual of Practice, Report No. FHWA-IF-99-004, Federal Highway Administration, U.S.

Department of Transportation.

3. Avent, R.R. and Mukai, D.J. (2001), "What you Should Know about Heat Straightening Repair of Damaged

Steel," Engineering Journal, 38(1), 27-49.

4. Avent, R.R., Mukai, D.J., and Robinson, P.F. (2000), “Heat Straightening Rolled Shapes,” Journal of Structural

Engineering, 126(7) 755-763.

5. Roeder, C.W. (1986). “Experimental Study of Heat Induced Deformation, ” Journal of Structural Engineering,

Vol. 112(10), pp. 2247-2262.

6. Avent, R.R., Mukai, D.J., Robinson, P.F., and Boudreaux, R.J. (2000), “Heat Straightening Damaged Steel

Plate Elements,” Journal of Structural Engineering, 126(7), 747-754.

7. Avent, R.R., Mukai, D.J., and Robinson, P.F. (2000), “Effect of Heat Straightening on Material Properties of

Steel,” Journal of Materials in Civil Engineering, 12(3), 188-195.

8. Avent, R. R., Mukai, D.J. and Robinson, P.F. (2001). “Residual-Stresses in Heat Straightened Steel Members,”

Journal of Materials in Civil Engineering, 13(1), 18-25.

9. Shanafelt, G.O., and Horn, W.G. (1984). Guidelines for Evaluation and Repair of Damaged Steel Bridge

Members, NCHRP Report No. 271, Transportation Research Board, National Research Council, Washington.

10. Avent, R.R. and Fadous, G.M. (1989). “Heat Straightening Prototype Damaged Bridge Girders,” Journal of

Structural Engineering, 115(7), 1631-1649.

11. Nichols, J. I. and Weerth, D. E. (1972). “Investigation of Triangular Heats Applied to Mild Steel Plates,”

Engineering Journal, AISC, Oct., 137-141.

12. Rothman, R. L. (1973), Flame Straightening Quenched and Tempered Steels in Ship Construction, Report No.

235, Ship Structures Committee.

13. Varma, A.H., and Kowalkowski, K.J. (2004). Effects of Multiple Damage-Heat Straightening Repair on the

Structural Properties of Bridge Steels, Report No. RC-1456, Michigan Department of Transportation, Lansing, .

14. Kowalkowski, K.J., and Varma, A.H. (2007). “Structural Properties of Steels Subjected to Multiple Cycles of

Damage Followed by Heating Repair.” Journal of Structural Engineering, ASCE, Vol. 133, No. 2, pp. 283-296.

15. Kowalkowski, K.J. (2005). “Effects of Multiple Damage-Heat Straightening Repair Cycles on the Structural

Properties and Serviceability of Steel Beam Bridges.” Ph.D. Dissertation, Purdue University, School of Civil

Engineering, West Lafayette, IN.

16. ASTM (2005). “Standard Test Methods for Tension Testing of Metallic Materials.” E8, American Socitey for

Testing and Materials, West Conshohocken, PA.

17. ASTM (2005). “Standard Test Methods for Rockwell Hardness and Rockwell Superficial Hardness of Metallic

Materials.” E18, American Society for Testing and Materials, West Conshohocken, PA.

18. ASTM (2005). “Standard Test Methods for Notched Bar Impact Testing of Metallic Materials.” E23, American

Society for Testing and Materials, West Conshohocken, PA.

19. AASHTO (2004). LRFD Bridge Design Specifications – 3rd

edition. Association of State Highway

Transportation Officials, Washington, D.C.

11

LIST OF TABLES

TABLE 1 Large-scale beam test matrix

TABLE 2 Undamaged steel material properties

TABLE 3 Normalized structural properties of damaged-repaired beam specimens

TABLE 4 Fracture toughness of damaged-repaired A7 beam specimens

TABLE 5 Fracture toughness results of damaged-repaired A588 beam specimens


LIST OF FIGURES

FIGURE 1 Bottom flange displacements of steel beams impacted by over-height trucks

FIGURE 2 Heat-straightening repair procedure used by MDOT Statewide Bridge Crew (SBC)

FIGURE 3 Test setup and instrumentation layout for Specimen A7-Beam 2: (a) before, and (b) after first damage

cycle

FIGURE 4 Schematic drawing of loading frame: (a) elevation, and (b) side view

FIGURE 5 Plastic strain distribution in A7-Beam-2 subjected to damage strain of 90!y

FIGURE 6 Vee heat locations and nomenclature (front side)

FIGURE 7 Midspan displacements during all three damage cycles of A36-Beam 1

FIGURE 8 Experimental behavior during the second repair cycle of A7-Beam 1

FIGURE 9 Original locations of material samples in the flanges of damaged-repaired beam specimens

12

TABLE 1 Large-scale beam test matrix

!d / !y Mr / Mp-y #p (mm) Tmax (°C)

Specimen ID Cycle No. Cycle No. Cycle No. Cycle No.

1 2 3 1 2 3 1 2 3 1, 2, 3

A7-Beam-1 30 30 30 0.25 0.50 0.25 56 56 56 650

A7-Beam-2 90 60 60 0.50 0.50 0.50 216 149 149 650

A588-Beam-1 40 20 20 0.25 0.25 0.25 125 53 53 650

A588-Beam-2 40 20 20 0.50 0.50 0.50 125 53 53 650

A36-Beam-1 30 30 30 0.25 0.50 0.25 79 79 79 650

A36-Beam-2 30 30 30 0.25 0.50 0.25 79 79 79 760

13

TABLE 2 Undamaged steel material properties

Steel/Plate "yo Eo "uo eo HRBo Fracture Toughness (J)

(MPa) (GPa) (MPa) (%) (B-Scale) Mean 95% H 95% L

A7 252.4 209.4 421.6 41.82 64.1 94.64 106.02 83.34

A588 405.1 207.0 585.0 35.66 87.0 129.75 138.16 121.32

A36 322.3 210.6 517.8 36.85 80.7 65.35 78.23 52.47

14

TABLE 3 Normalized structural properties of damaged-repaired beam specimens

Specimen Coupon "y / "yo E / Eo "u / "uo e / eo HRB / HRBo

X 1.15 1.00 1.01 0.96

Y 1.16 0.94 1.02 0.96 A7-Beam-1

Z 1.25 0.93 1.06 0.81

1.07

X 1.07 0.95 1.03 0.80

Y 1.15 0.96 1.02 0.94 A7-Beam-2

Z 1.20 0.94 1.03 0.80

1.12

X 0.94 1.00 0.90 1.03

Y 0.96 0.99 0.92 0.84 A588-Beam-1

Z 1.07 1.01 0.98 0.79

0.95

X 0.90 0.97 0.90 1.02

Y 0.94 1.02 0.90 0.93 A588-Beam-2

Z 1.08 0.98 0.99 0.81

0.94

X 1.07 1.00 0.98 0.97

Y 1.07 0.93 0.96 0.94 A36-Beam-1

Z 1.07 1.04 0.98 0.88

1.10

X 1.21 1.04 1.05 0.89

Y 1.16 0.97 0.99 0.69 A36-Beam-2

Z 1.10 0.98 0.98 0.79

0.97

15


A7-Beam-1 A7-Beam-2

L1 C R1 Avg-H L1 C R1 Avg.-H

1 69 12 24 35 191 100 268 187

2 22 7 14 14 228 79 84 130

3 18 14 12 15 98 11 83 64

4 11 9 11 11 12 20 17 16

Avg-V 30 11 15 19 133 53 113 99

Normalized fracture toughness (i.e., divided by 94.64 J)

L1 C R1 Avg-H L1 C R1 Avg.-H

1 0.73 0.13 0.25 0.37 2.03 1.07 2.85 1.98

2 0.23 0.07 0.14 0.15 2.42 0.84 0.89 1.38

3 0.19 0.14 0.13 0.15 1.04 0.12 0.88 0.68

4 0.12 0.09 0.12 0.11 0.13 0.22 0.18 0.18

Avg-V 0.32 0.11 0.16 0.20 1.40 0.56 1.20 1.05

16


A588-Beam-1 A588-Beam-2

L1 C R1 Avg-H L1 C R1 Avg-H

1 400 400 400 400 400 397 390 396

2 399 396 399 397 397 373 353 374

3 359 342 396 366 176 163 144 161

4 191 164 176 178 141 100 138 126

Avg-V 338 325 343 335 279 259 256 264



1 3.08 3.08 3.08 3.08 3.08 3.06 3.01 3.05

2 3.07 3.05 3.07 3.07 3.06 2.87 2.72 2.88

3 2.77 2.63 3.05 2.82 1.36 1.25 1.11 1.24

4 1.47 1.26 1.36 1.37 1.09 0.77 1.07 0.98

Avg-V 2.60 2.51 2.64 2.58 2.15 1.99 1.97 2.04

17


A36-Beam 1 A36-Beam 2


1 92 22 132 81 125 278 244 216

2 89 92 50 77 152 193 255 199

3 23 22 87 43 125 75 146 115

4 14 15 22 18 61 24 99 61

Avg-V 54 38 73 56 115 142 186 148



1 1.41 0.33 2.01 1.25 1.91 4.25 3.73 3.30

2 1.37 1.41 0.77 1.18 2.32 2.95 3.90 3.06

3 0.35 0.33 1.33 0.67 1.91 1.14 2.24 1.76

4 0.22 0.23 0.33 0.26 0.93 0.37 1.51 0.94

Avg-V 0.84 0.58 1.11 0.84 1.77 2.28 2.85 2.27

18

FIGURE 1 Bottom flange displacements of steel beams impacted by over-height trucks

19

FIGURE 2 Heat-straightening repair procedure used by MDOT Statewide Bridge Crew (SBC)

20

FIGURE 3 Test setup and instrumentation layout for Specimen A7-Beam 2: (a) before damage, and (b) after

first damage cycle

21

FIGURE 4 Schematic drawing of loading frame: (a) elevation, and (b) side view

22

FIGURE 5 Plastic strain distribution in A7-Beam-2 subjected to damage strain of 90!y

23

FIGURE 6 Vee heat locations and nomenclature (front side)

24

FIGURE 7 Midspan displacements during all three damage cycles of A36-Beam 1

25

FIGURE 8 Experimental behavior during the second repair cycle of A7-Beam 1

26

FIGURE 9 Original locations of material samples in the flanges of damaged-repaired beam specimens

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