what you should know about heat straightening repair of damaged steel

ENGINEERING JOURNAL / FIRST QUARTER / 2001 / 27

INTRODUCTION

The repair of damaged steel through the application ofheat has long been considered an art rather than a sci-

ence. As such both the methodology and results have var-ied considerably from one practitioner to another. Becausesteel tends to be a forgiving material, successful repairs out-number failures. Yet material properties may be compro-mised resulting in a weakened in-place structure. Inaddition, it is not unusual for a fracture to occur during therepair process. As a result, the use of heat straightening (orflame straightening) has had limited application. Manyengineers are reluctant to allow heat straightening becauseof the lack of an established method of application as wellas the lack of quantification of steel properties after therepair is complete. The purpose of this paper is to providesome engineering guidelines for use of the heat-straighten-ing process.

The ideas presented in this paper are referenced to sup-porting research and provide the reader with a place tobegin when contemplating heat straightening repair. Thepaper is intended to be a state-of-the-practice report, whichshould be used to complement sound engineering judgment,not to replace it. In addition, many of the principles dis-cussed here also apply to heat curving to produce camber orsweep.

WHAT IS HEAT SRAIGHTENING?

Heat-straightening is a repair procedure in which a limitedamount of heat is applied in specific patterns to the plasti-cally deformed regions of damaged steel in repetitive heat-ing and cooling cycles to produce a gradual straightening ofthe material. The process relies on internal and externalrestraints that produce thickening (or upsetting) during theheating phase and in-plane contraction during the coolingphase. Heat straightening is distinguished from other meth-ods in that force is not used as the primary instrument of

straightening. Rather, the thermal expansion/contraction isan unsymmetrical process in which each cycle leads to agradual straightening trend. The process is characterized bythe following conditions, which must be maintained:

1. The maximum heating temperature of the steel doesnot exceed either (a) the lower critical temperature(the lowest temperature at which permanent molecu-lar changes may occur), or (b) the temper limit forquenched and tempered steels.

2. The stresses produced by applied external forces donot exceed the yield stress of the steel in its heatedcondition.

3. Only the regions in the vicinity of the plasticallydeformed zones are heated.

When these conditions are met, the material propertiesundergo relatively small changes and the performance ofthe steel remains essentially unchanged after heat straight-ening. Properly conducted, heat straightening is a safe andeconomical procedure for repairing damaged steel.

HOW DOES HEAT SRAIGHTENING WORK?

The basic concept of heat straightening is relatively simpleand relies on two distinct properties of steel:

• If steel is stretched or compressed past yield, it doesnot assume its original shape when released. Rather,it remains partially elongated or shortened, dependingon the direction of the originally applied force.

• If steel is heated to relatively modest temperatures(370-700°C or 700-1300°F), its yield value becomessignificantly lower while at the elevated temperature.

To illustrate how steel can be permanently deformedusing these two properties, consider the short steel bar inFigure 1a. First, the bar is placed in a fixture, muchstronger than the bar itself, and clamped snug-tight (Figure1b). Then the bar is heated in the shaded portion. As thebar is heated it tries to expand. However, the fixture pre-vents expansion in the longitudinal direction. Thus, the fix-ture exerts restraining forces on the bar as shown in Figure1c. Since the bar is prevented from longitudinal expansion,it is forced to expand a greater amount laterally and trans-versely through its thickness than in an identical unre-strained bar. Consequently, a bulge will occur in the heatedzone. Because the bulge has been heated, its yield value hasbeen lowered, resulting in plastic deformations, which do

What You Should Know About Heat Straightening Repair of Damaged SteelR. RICHARD AVENT and DAVID J. MUKAI

R. Richard Avent is Robert and Adele Anding distinguishedprofessor, department of civil and environmental engineer-ing, Louisiana State University, Baton Rouge, LA.

David J. Mukai is assistant professor, department of civil andenvironmental engineering, Louisiana State University,BatonRouge, LA.

28 / ENGINEERING JOURNAL / FIRST QUARTER / 2001

not occur in the unheated portions. When the heatingsource is removed, the material will cool and contract three-dimensionally. The clamp cannot prevent the bar from con-tracting longitudinally. As cooling progresses the barshortens and the bulge shrinks. However, a portion of thebulge is permanent even after the bar has completely cooledand the bar is shorter than its original length, Figure 2.1d.In essence a permanent redistribution of material hasoccurred in the heated zone leaving the bar slightly shorterwith a small bulge. This permanent bulge, or thickening, inthe heated zone is called “upsetting”. The redistribution ofmaterial is referred to as “plastic deformation” or “plasticflow”. The clamping force is often referred to as a restrain-ing force. Through cycles of clamping, heating, and cool-ing, the bar could be shortened to practically any lengthdesired.

This simple example illustrates the fundamental princi-ples of heat straightening. However, most damage in steelmembers is much more complex than stretching or shorten-ing of a bar. Consequently, different damage conditionsrequire their own unique heating and restraining patterns.

WHY PLACE LIMITS ON HEATINGTEMPERATURE AND JACKING STRESSES?

A clear distinction should be made between heat straighten-ing and two other methods often confused with heatstraightening: hot mechanical straightening and hot work-ing. Hot mechanical straightening differs from heat

straightening in that applied external force is used tostraighten the damage. These applied forces producestresses well above yield, resulting in large movements dur-ing a single heat cycle. Often the member is completelystraightened by the continued application of a large forceduring a single cycle. The results of this type of straighten-ing are unpredictable and little research has been conductedon this procedure. Specific concerns about hot mechanicalstraightening include:

1. Fracture may occur during straightening2. Material properties may be adversely affected3. Buckles, wrinkles or crimps may resultThe Engineer should recognize that hot mechanical

straightening is an unproven method, which may lead todamaged or degraded steel. As such, its use should be con-sidered only in special cases in which other methods are notviable.

Hot working is distinguished from heat straightening inthat both large external forces and high heat are used. Thismethod is similar to hot mechanical straightening in thatexternal forces are used. In addition, the steel is heated wellabove the lower critical temperature and often glows cherryred indicating a temperature above the upper critical tem-perature. The results of this process are highly unpre-dictable and may result in:

1. Fracture during straightening2. Severe changes in molecular structure which may not

be reversible3. Severe changes in mechanical properties including a

high degree of brittleness4. Buckles, wrinkles, crimps, and other distortionsHot working should not be used to repair damaged steel.Some practitioners will tend to over-jack and over-heat

yet claim to be heat straightening. The reader is cautionedto be aware of these distinctions when specifying heatstraightening as opposed to either hot mechanical straight-ening or hot working.

WHAT TEMPERATURE LIMITS ARE NECESSARY WHEN HEAT STRAIGHTENING?

One of the most important and yet difficult-to-controlparameters of heat straightening is the temperature of theheated metal. Factors affecting the temperature include sizeand type of the torch orifice, intensity of the flame, speed oftorch movement, and thickness and configuration of themember. Assuming that adequate control of the appliedtemperature is maintained, the question arises as to whattemperature produces the best results in heat straighteningwithout altering the material properties.

The large majority of steels used for construction in theUnited States are either carbon or low alloy steel. At ambi-ent temperature, these steels have two major constituents:

Fig. 1. Conceptual example of shortening a steel bar.


ferrite grains and pearlite grains. Pearlite is itself composedof alternating laths of ferrite iron and cementite (Fe3C). Aniron-carbon equilibrium diagram is shown in Figure 2 illus-trating the relationship between the components of steel andtemperature. Solid steel is comprised of grains. The grainsthemselves are composed of iron atoms arranged on a crys-tal lattice. The carbon (and for that matter any other alloyedor residual atoms) do not molecularly bond with the iron,but rather; nest within the spaces between the iron atoms(interstitial solid solution), or displaces an iron atom from alattice node and occupies the position itself (substitutionalsolid solution). Ferrite is essentially pure iron. However,by examining the iron-carbon Phase diagram (Figure 2), itcan be seen that ferrite has a carbon solubility capacity upto .008 percent. Cementite is an orthorhombic crystal struc-ture comprised of iron and carbon. Chemically, cementitecontains 6.67 percent carbon dissolved in the crystal matrix(but is not molecularly bonded to the iron). The ratio of fer-rite to cementite in pearlite formed under slow cooled andunder equilibrium conditions from a eutectoid steel isapproximately 6:1. However, under other conditions (slowcool as not to form martensite or bainite) and compositions,pearlite has been observed to contain ferrite in a 3:1 to 2:1ratio with cementite. A low carbon steel has less than 0.8percent carbon, too little carbon to develop a 100 percentpearlite structure, resulting in pearlite plus ferrite. This fer-rite can exist as both a grain and as a constituent of pearlite.Carbon plus iron alloys with less than 0.8 percent carbonare hypoeutectic steels, commonly known as ferritic steels.These steels have a larger volume fraction of ferrite grains,body-centered crystal structure (BCC), than pearlite. Highcarbon steels (carbon content between 0.8 and 2.0 percent)are hypereutectic steels, commonly known as pearlitesteels. These steels are predominately pearlite (orthorhom-bic cementite and BCC ferrite). The volume fraction ofcementite and ferrite are equal at the eutectic composition

(0.8 percent carbon). The volume fraction of cementiteincreases with increasing carbon content of steel. Low car-bon steels tend to be softer and more ductile because theseare characteristics of ferrite. Cementite is hard and brittlethus high carbon steels are harder and less ductile.

Temperatures greater than 727°C (1340°F) begin to pro-duce a phase change in steel. This temperature is oftencalled the lower critical (or lower phase transition) temper-ature. The ferritic and pearlite crystal structures begin toassume a face centered cubic form (austenite). Thus, as fer-rite and pearlite are heated to the austenitic range and theircrystal lattices change to face-centered cubic, all containedcarbon goes into solution. When steel cools below thelower critical temperature, it attempts to return to its formerferritic or pearlitic structure. Since this change requires aspecified time frame, rapid cooling may not permit thecomplete molecular change to occur. Under these circum-stances, a hard, strong and brittle phase called martensiteoccurs. The steel in this form may have reduced ductilityand be more sensitive to brittle fracture under repeatedloads.

The upper critical (or upper phase transition) temperatureis the level at which the molecular change in structure iscomplete. At this temperature (around 815 to 925°C or1500 to 1700°F for most steels, depending on carbon con-tent) the steel assumes the form of a uniform solid solutioncalled austenite. It is at temperatures between the lower andupper critical that a wide range of mill hot rolling and work-ing can occur. As long as the temperature is lowered slowlyin a controlled manner from these levels, the steel assumesits original molecular configuration and properties. Thistemperature control is more difficult to maintain at a fabri-cation shop or in the field when conducting heat straighten-ing repairs. Consequently, if the temperature during heatstraightening is not kept below the lower critical tempera-ture, undesirable properties may be produced during cool-ing.

It has been shown by Avent, Fadous, and Boudreaux(1991), Avent and Mukai (1998), and Roeder (1986), thatthe amount of movement during heat straightening of dam-aged flexural members is directly proportional to heatingtemperature up to at least 870°C (1800°F). Permanentmovement can occur with heating temperatures as low as370°C (700°F). To avoid the possibility of detrimentalchanges in material properties, it is recommended that thesteel temperature be kept somewhat below the lower phasetransition temperature of approximately 700°C (1300°F).

The maximum temperature recommended by mostresearchers such as Holt (1971), Avent (1995), Roeder(1986) and Shannafelt and Horn (1984), is 650°C (1200°F)for all but quenched and tempered high-strength steels. Thelimiting temperature of 650°C (1200°F) allows for about55°C (100°F) of temperature variation, which was found toFig. 2. Iron-carbon equilibrium diagram.

be a common range among experienced practitioners. TheAmerican Welding Society Bridge Welding Code (1996)specifies maximum heating temperatures of 590°C(1100°F) for quenched and tempered steels and 650°C(1200°F) for all others. For A514 and A709 (grades 100and 100W), a minimum tempering temperature of 620°C(1150°F) is required. Thus, the 590°C (1100°F) limit pro-vides a 30°C (50°F) safety factor. However, for A709 grade70W the specified minimum tempering temperature is590°C (1100°F). A maximum heating temperature of565°C (1050°F) is recommended for this grade to provide a30°C (50°F) safety factor and to avoid property changes.

To control the temperature, the speed of the torch move-ment and the size of the orifice must be adjusted for differ-ent thicknesses of material. However, as long as thetemperature is rapidly achieved at the appropriate level, thecontraction effect will be similar. Various methods can beused to monitor temperature during heating. Principalamong these includes: visual observation of color of thesteel; use of special temperature sensing crayons or pyrom-eters; and infrared electronic temperature sensing devices.

HOW SHOULD THE STEEL BE COOLED AFTER HEATING?

To obtain the maximum effect of heat straightening, thesteel should be allowed to cool to approximately ambienttemperature. If not, the desired temperature gradient anddistribution during the succeeding heating cycle will bechanged to produce a less effective pattern. It is recom-mended that the steel cool to at least 90°C (200°F) prior toapplying the next cycle.

The method of cooling most often used is natural coolingwithout artificial aids. However, this approach slows therepair and increases costs. Theoretically, any artificial cool-ing procedure could be used as long as the maximum tem-perature in the steel does not exceed the lower phasetransition temperature or the lowest tempering temperaturesince changes in crystal structure would not occur. From apractical viewpoint, it is difficult to ensure that heating tem-perature can be perfectly controlled. Therefore, it is rec-ommended that the steel be allowed to cool naturally to315°C (600°F) prior to using compressed air or water toaccelerate cooling.

WHAT LIMITS SHOULD BE PLACED ON THE MAGNITUDE OF JACKING FORCES?

On several occasions, the writers have observed suddenfractures during heat straightening. These types of fracturesare not uncommon. The usual explanation is that smallmicrocracks must have occurred as a result of initial dam-age. Thus, during heat straightening, one of these crackspropagates. However, no experimental evidence validating

this assumption has been published. The senior writer hasinvestigated this phenomenon (Avent and Fadous, 1989).Successively higher jacking forces were used in heatstraightening minor axis damage in a W24×76 beam. Thisbeam suddenly fractured during the eighth heating cyclewith the highest jacking force applied. Based on the move-ment per heat as well as jacking force measurements, thejacking stresses exceeded the reduced yield (but notreduced tensile strength) in the heated zones. It was con-cluded that over-jacking, which produced stresses that wereadditive to residual stresses, produced this failure. Typi-cally, heat straightening practitioners do not use calibratedjacks but rather control the force by “feel”. The number ofvariables makes this approach likely to produce over-stress.For example, consider a composite bridge girder. The fol-lowing variables are associated with inducing stress byjacking the bottom flange: diaphragm spacing, number ofdiaphragms, beam length, section modulus of beam, depth-to-web thickness ratio, and section modulus of lowerflange. For a given stress level, the interaction of so manyvariables means that the jacking force could vary over awide range. By using the “feel” method, it is erroneouslyassumed that the jacking force varies little from one struc-ture to the next. Thus, it would not be unusual to over-stresssuch a beam with the likely result of a fracture during heatstraightening.

It is therefore recommended that stresses in the heatedzone due to jacking not exceed 50 percent of the nominalyield stress at ambient temperature. Since the yield stressat 1200°F is reduced by approximately 50 percent, this cri-terion prevents the heated steel stress from becoming sig-nificantly larger than yield. In addition, because the entirecross-section does not reach the maximum heating temper-ature simultaneously, the 50 percent value includes an addi-tional margin of safety. Thus, it is also recommended thatjacks always be calibrated and the structure be analyzed toprevent over-stress.

HOW SHOULD HEAT BE APPLIED TO THE STEEL?

The primary equipment utilized for heat straightening is aheating torch. The heat source is typically an oxygen-fuelmixture. Typical fuels include acetylene, propane, and nat-ural gas. The appropriate fuel is mixed with oxygen underpressure at the nozzle to produce a proper heating flame.Either a single or a multiple orifice tip may be used. Thesize and type is dictated by the fuel selected and thicknessof material to be heated. A No. 8 single orifice tip is gener-ally satisfactory for thicknesses up to 20 to 25 mm (3/4 or 1in.) with acetylene. For thinner material a smaller tip is rec-ommended. If heavy sections are being heated, a single ori-fice tip may not be adequate. For such cases, either heat


from both sides or use a rosebud or multiple orifice tip. Thesize may vary depending on the material thickness. Thedetermining factor is the ability to raise the through-thick-ness steel temperature to the specified level. Note thatwhether a single or multiple orifice is used, the torch shouldbe a heating torch and not a cutting torch.

WHAT ARE THE BASIC TYPES OF HEATS?

There are several basic types of heats that are frequentlyused as illustrated by Avent et al. (1991), Avent and Mukai(1998), Ciesicki and Bulter (1968), and Holt (1955, 1965,1971, 1977). For some applications only one type isrequired. However, in many cases a combination of thesebasic types is necessary. The three basic types most oftenused in heat straightening are described as follows.

Vee Heat

The vee heat is used to straighten strong axis bends in steelplate elements. As seen in Figure 3, a typical vee heat startsat the apex of the vee-shaped area using an oxy-fuel torch.When the desired temperature is reached (usually around650°C or 1200°F for mild carbon steel), the torch isadvanced progressively in a serpentine motion toward thebase of the vee. This motion is efficient for progressivelyheating the vee from top to bottom. The plate will initiallymove upward (Figure 3a) as a result of longitudinal expan-sion of material above the neutral axis producing negativebending. The cool material adjacent to the heated area

resists the normal thermal expansion of the steel in the lon-gitudinal direction. As a result, the heated material willtend to expand, or upset, to a greater extent through thethickness of the plate than longitudinally, resulting in plas-tic flow. At the completion of the heat, the entire heatedarea is at a high and relatively uniform temperature. At thispoint the plate has moved downward (Figure 3b) due to lon-gitudinal expansion of material below the neutral axis, pro-ducing positive bending. As the steel cools, the materialcontracts longitudinally to a greater degree than the expan-sion during heating. Thus, a net contraction occurs.Because the net upsetting is proportional to the width acrossthe vee, the amount of upsetting increases from top to bot-tom of the vee. This variation produces a closure of the vee.Bending is produced in an initially straight member, orstraightening occurs (if the plate is bent in the oppositedirection to that of the straightening movement, Figure 3c).For many applications, it is most efficient to utilize a veethat extends over the full depth of the plate element, butpartial depth vees may be applicable in certain situations.When using partial depth vees, the open end should extendto the edge of the element. The vee depth is varied by plac-ing the apex at a partial depth location. For example, indeep plate elements the depth of the vee may result in toomuch cooling near the apex prior to completing the heat atthe open end. A 3/4 depth vee may be applicable in reducingthis problem. Other examples of partial depth vees will beshown in later sections.

Line Heats

Line heats are employed to repair a bend in a plate about itsweak axis. Such bends, severe enough to produce yieldingof the material, often result in yield lines. A line heat con-sists of a single pass of the torch. The restraint in this caseis often provided by an external force although movementwill occur without external constraints. This behavior isillustrated in Figure 4. A line heat is applied to the under-side of a plate element subjected to bending moments pro-duced by external forces (Figure 4a). As the torch isapplied and moved across the plate, the temperature distri-bution decreases through the thickness (Figure 4b). Thecool material ahead of the torch constrains thermal expan-sion, even if bending moment constraints are not present.Because of the thermal gradient, more upsetting occurs onthe torch (or hotter) side of the plate. During cooling thisside consequently contracts more, creating a concave bendon the torch side of the plate similar to that shown in Figure4d. Thus, to straighten a plate bent about its weak axis, theheat should be applied to the convex side of the damagedplate. The movement can be magnified by the use ofapplied forces, which produce bending moments about theyield line (Figure 4c). Referring to a section through the


Fig. 3. Stages of movement during vee heat.

plate transverse to the line heat (Figure 4c), the restrainingmoments tend to prevent transverse expansion below theplate centerline. In a manner similar to the vee heat mech-anism, the material thus tends to expand through the thick-ness, or “upset”. Upon cooling, the restraining momentstend to magnify transverse contraction (Figure 4d). Thespeed of the travel of the torch is critical as it determines thetemperature attained. With proper restraints and a uniformspeed of the torch, a rotation will occur about the heatedline.

Strip Heats

Strip heats, also called rectangular heats, are used to com-plement a vee heat. Strip heats are similar to vee heats andare accomplished in a like manner. Beginning at the initia-tion point, the torch is moved back and forth in a serpentine

fashion across a strip for a desired length, Figure 5a. Thispattern sequentially brings the entire strip to the desiredtemperature. The orientation can be an important consider-ation. The strip heat may be initiated at the midpoint andmoved toward both edges simultaneously using twotorches. A second alternative with similar effect is shownin Figure 5b using a single torch and starting from one side.Depending on the structural configuration, the strip mayalso be started at a free edge as shown in Figure 5c. How-ever, without restraints, this orientation may produce someweak axis bending. By alternating the initiation point toopposite edges in successive heating cycles, the weak axisbending can be minimized.

IS THERE A RELATIONSHIP BETWEEN FUNDAMENTAL DAMAGE PATTERNS AND

BASIC TYPES OF HEATS?

A convenient way to classify damage is to define four fun-damental damage patterns (Avent, 1995). For each damagepattern, there is a sequence of specific heating types thatwill produce straightening. In most cases, actual damage tosteel structures is a combination of two or more of thesefundamental damage patterns. However, more complexdamage can be repaired by sequentially using the heatingpatterns for these four fundamental damage patterns. Forpurposes of defining heating patterns, it is convenient torefer to the elements of a cross section as either primary orstiffening elements. The primary elements are plate ele-ments damaged by bending about their major axes. Thestiffening elements are those bent about their minor axes.Typically, vee heats are applied to primary elements whilestrip, line or no heat at all may be applied to stiffening ele-ments. The combination of these heats, applied eithersimultaneously or consecutively, represent a single cycle ofthe heating pattern. Because the net change in curvatureafter one heating cycle is small, a number of cycles arerequired to straighten a damaged member. For each cycle


Fig. 4. Schematic of line heat mechanism. Fig. 5. Schematic of strip heat on the top flange of a wide flange beam.

the heating pattern should be shifted to a different locationwithin the yield zone region, although the same location canbe returned to after every third cycle. More heats should beplaced in the region of largest curvature and fewer nearextremities to reflect the difference in damage curvature.The fundamental damage categories and their heating pat-terns are described in the following sections.

Category S

This type refers to bending about the “strong” or major axis.The heating pattern consists of a vee and strip heat as shownin Figure 6. This pattern is superimposed on a typical yieldzone for a wide flange beam bent about the strong axis. Thestrip heat and open end of the vee are placed on the flangedamaged in tension. The vee heat is first applied to theweb. Upon completion, a strip heat is applied to the flangeat the open end of the vee. The width of the strip heatalways equals the vee width at the point of intersection.This procedure allows the vee to close during cooling with-out restraint from the stiffening element. No heat is appliedto the flange at the apex of the vee. This vee/strip combi-nation is repeated by shifting over the vicinity of the yieldzone until the member is straight. The pattern is shown fora wide flange and channel in Figure 6.

Category W

This category refers to damage as a result of bending aboutthe “weak” or minor axis. For rolled or built-up shapes theweb is usually at, or near, the neutral axis. Consequently, itmay not deform into the inelastic range. The heating pat-tern for this case is similar to the previous case but note theprimary and stiffening elements are reversed. The vee heatis first applied to both flanges (either simultaneously or one

at a time) as shown in Figure 7. After heating these primaryelements, a strip heat is applied to the web. The only excep-tion is that no strip heat is applied to stiffening elementslocated adjacent to the apex of a vee heated element sincethis element offers little restraint to the closing of the veeduring cooling. Note that the width of the strip heat is equalto the width of the vee heat at the point of intersection. Forall cases the pattern is repeated by shifting over the vicinityof the yield zone until the member is straight.

Category T

This type refers to damage as a result of torsion or twistingabout the longitudinal axis of a member. For rolled or built-up shapes, the flange elements tend to exhibit flexural plas-tic deformation in opposite directions. The web is oftenstressed at levels below yield. If one flange is constrained(such as the case of a composite beam), then only theunconstrained flange element is subjected to plastic defor-mation and yielding may also occur in the web. The heat-ing pattern for this damage case is shown in Figure 8. Thevees on the top and bottom flange are reversed to reflect thedifferent directions of curvature of the opposite flanges.The vee heats are applied first and then the strip heat isapplied. Note that for the channel, the strip heat need onlybe applied to half depth. This half depth strip allows thelower flange vee to close with minimal restraint from theweb. If one flange is a composite connection, then thatflange is not heated.


Fig. 6. Heating patterns for wide flanges and channels bent about theirmajor axes (Category S).

Fig. 7. Heating patterns for wide flanges and channels bent about theirminor axes (Category W).

Category L

This category includes damage that is localized in nature.Local flange or web buckles, web crippling and small bendsor crimps in plate elements of a cross section typify thisbehavior. A local buckle or bulge reflects an elongation ofmaterial. Restoration requires the bulging area to be short-ened. A series of vee or line heats can be used for this pur-pose as shown in Figure 9. These vees are heatedsequentially across the buckle or around the bulge. For webbulges either lines or vees may be used. If vees are used,they are spaced so that the open end of the vees touch.There is a tendency for practitioners to over-heat webbulges. For most cases, too much heat is counter-produc-tive. The preferred pattern is the line heats in thespoke/wagon wheel pattern. For the flange buckle pattern(Figure 9b), either lines or a combination of lines and veesmay be used. For most cases, the line pattern with few orno vees tends to be most effective. Since the flange dam-age tends to be unsymmetrical, more heating cycles arerequired on the side with the most damage. When possible,the heating should be done on the convex side.

IS THE SEQUENCING OF HEATING IMPORTANT?

When a combination of vee, strip or line heats is used, theorder of heating is referred to as the sequence. Thesequencing of heats may be important in some straighteningoperations. However, little research has been conducted toverify its effects. Some practitioners feel that propersequencing will accelerate the straightening and help keepresidual stresses to a minimum. Consider the case of an I-beam with Category S damage requiring a vee heat in theweb and a strip heat in the flange as shown in Figure 6.

A common sequence is to heat the vee first, followedimmediately by the strip. An alternative used by some prac-titioners is to heat the web vee first and allow it to cool fora few minutes before heating the flange strip. It is felt bysome that this approach will reduce the residual stresses sig-nificantly. The available research data and differentsequences used in practice indicate that more than onesequence can be successful for many cases. Since differentsequences are successfully used in practice, it appears thatheat straightening is not overly sensitive to the sequenceused. At this time there is not adequate documentation tosupport one sequence over another for a particular heatingpattern. The experience of the practitioner is the most reli-able guide to proper sequencing. The sequencing patternsshown here are based on those often successfully used inpractice.

HOW CAN THE HEATING PATTERNS FOR THEFOUR FUNDAMENTAL DAMAGE TYPES BEMODIFIED FOR MORE GENERAL DAMAGE

CONDITIONS?

Since typical damage is often a combination of these fun-damental damage categories, a combination of heating pat-terns is often required. The key is to select the combinationof patterns to fit the damage. When in doubt, a good policyis to address one of the fundamental damage patterns at atime. For example, remove the Category W damage prior


Fig. 8. Wide flanges and channels with twisting damage (Category T). Fig. 9. Typical heating patterns for local damage.

to addressing the Category L damage. It should be notedthat with proper combinations, several types of damagecould be removed expeditiously. For example, suppose thata wide flange section is impacted such that the bendingoccurs about an axis at an arbitrary angle to the principalaxes, i.e., bending occurs about both the strong and weakaxis. The heating pattern, Figure 10, requires a vee heat onthe web to restore the strong axis damage and vee heats onthe flanges to restore the weak axis damage. The heatsshould be executed sequentially as numbered in Figure 10.Note that no strip heat is required on the web since a vee isused there. Restraining forces should be used to producebending moments about both the strong and weak axis asindicated in Figure 10, tending to straighten the damage.Once the damage is corrected about one of the principalaxes, the heating pattern should revert to one of the funda-mental patterns until straightening is complete about theother principal axis.

As a second example, consider a wide flange beam withweak axis bending damage combined with a local bulge inone flange. The heating pattern is shown in Figure 11. Veeheats are used on the top and bottom flanges along with aweb strip heat similar to the standard weak axis pattern.However, partial depth vees are used on the flange with thebulge along with a series of line heats along bulge yieldlines. Since a yield line is likely to occur at the lower webfillet, a line heat is also needed on the web. Restrainingforces are used to create bending moments about the weakaxis as shown in Figure 11. In addition, a jacking forceshould be applied on the local bulge as shown on the crosssection in Figure 11. The sequence of heats is also indi-cated in the figure.


A final example is the case of multiple plastic hingesformed about the weak axis such as might occur for a beamcontinuous over interior supports. The heating pattern isshown in Figure 12. Note the reversed direction of the veesto reflect the multiple curvature damage. The restrainingmoments must also reflect the reverse curvature nature ofthe damage as shown in the figure.

WHEN EVALUATING DAMAGE, HOW SHOULD ITBE QUANTIFIED?

It is important to quantify the level of damage prior todeciding if heat straightening is the best repair approach.Of particular interest is the degree to which the material hasbeen strained past yield. A useful measure is the strainratio, µ, which is the ratio of the maximum actual strain tothe strain at initial yield. The maximum actual strain, εmax,can be expressed in terms of radius of curvature, R, as

where ymax is the distance from the extreme fiber to the neu-tral axis. The yield strain, εy, is

where Fy is the yield stress and E is the modulus of elastic-ity. Combining these equations, the strain ratio is

εmax max= 1

Ry (1)

εyyF

E= (2)

µε

=y

RFy

max (3)

Fig. 10. Heating pattern and sequence for bending combination aboutboth the strong and weak axis.

Fig. 11. Heating pattern and sequence for combination weak axis bend-ing and local flange bulge.

If the material properties are known, the strain ratio canbe computed once the radius of curvature has been deter-mined. One convenient approach is to determine the radiusof curvature by measuring the degree of damage. Thedegree of damage, ϕd is defined as the change in slope of anelement over the yield zone as shown in Figure 13. ϕd canbe computed as:

where ϕd is the degree of damage or angle of permanentdeformation at the plastic hinge and yi is a measured offsetas shown in Figure 13b. If C is defined as the chord lengthbetween tangents at the edges of the yield zone, then

A more accurate approach is to take offsets within thedamaged zone where curvature is the sharpest. From Fig-ure 13a the radius of curvature, R, can be directly approxi-mated as

Using either approach the strain ratio can be calculatedfrom Equation 3 once R has been determined.

ARE THERE LIMITS ON THE LEVEL OFDAMAGE THAT CAN BE HEAT STRAIGHTENED?

Surprisingly little information is available on the effect ofdamage level (strain history). It is known that cold bendinginto the yield range reduces the ductility of steel in general

(Brockenbrough and Johnston, 1968). However, the appli-cation of heat tends to restore the original material charac-teristics. Shanafelt and Horn (1984) recommended that themaximum allowable strain be limited to 15 times the yieldstrain and/or 5 percent nominal strain for repair of tensionmembers. The limit approximately defines the delineationbetween the plastic region and the strain-hardening region.However, these recommendations were not backed by spe-cific research data. No limits were suggested for compres-sion members. A 5 percent nominal strain corresponds to aµ = 15 for Fy = 100 ksi. The writers have conducted a num-ber of tests on plates and beams in which the maximumstrain ratio, µ, was 100. Each of these members was suc-cessfully heat straightened. Stress-strain curves taken fromtensile coupons after heat straightening indicated that therewas no significant degradation of the material. These testsindicate that the limiting strain for heat straightening repairshould be equal to or greater than 100 times the yield stain.Based on anecdotal evidence even larger strains can proba-bly be repaired. However, research has not been conductedpast 100 times the yield strain.

WHAT IS THE IMPACT OFHEAT STRAIGHTENING REPAIR ON YIELD

STRESS AND TENSILE STRENGTH?

Most testing for the basic mechanical properties of heat-straightened plates and beams has been conducted onundamaged material after 3 or 4 vee heats (Welding Engi-neer, 1959; Nicholls and Weerth, 1972; Pattee, Evans, andMonroe, 1969, 1970; Roeder, 1986; Rothman, 1973; andRothman and Monroe, 1973). Researchers concluded fromthese tests that: (1) little change occurred in modulus of


ϕdy y

L

y y

L= − + −− −tan ( ) tan ( )1 2 1

1

1 3 4

2

(4)

1 2

2R Cd= ( )sin

ϕ(5)

1 21 12R

y y y

Lr r r=

− +− + (6)

Fig. 12. Heating pattern for reverse curvature bending. Fig. 13. Offset measurements to calculate degree of damage and radiusof curvature.

elasticity, (2) slight increases were found in yield and ulti-mate tensile stress, and (3) 10 to 25 percent reduction inductility was observed. Of more significance are the prop-erties of damaged plates (or rolled shapes) after experienc-ing the large number of heats required to fully straighten themember. Avent, Robinson, Madan, and Shenoy (1993a),Avent and Mukai (1998) conducted tests on seven platesand four beams which were damaged to strain ratios rang-ing from 30 to 100 and then heat straightened. In general,it was found that the yield stress in the heated regionsincreased approximately 10 percent over that of theunheated member. The largest increases were found at theapex of the vee heats where yield increased by 15 to 20 per-cent. The maximum tensile stress values showed a smallerincrease of around 4 to 6 percent. This test data indicatesthat heat straightening has a small but significant heat-treat-ing effect on steel.

WHAT IS THE IMPACT OFHEAT STRAIGHTENING ON

MODULUS OF ELASTICITY AND DUCTILTY?

Tests on undamaged plate specimens have suggested thatheat straightening had little effect on modulus of elasticity.However, the writers’ tests, Avent et al. (1993a), on actualdamaged and repaired plates and beams indicated reduc-tions on the order of 8 to 23 percent. The most significantchanges occurred in ductility, which was typically reducedby one-third. While this loss of ductility is significant, allmembers tested still met ASTM specifications.

WHAT IS THE IMPACT OFHEAT STRAIGHTENING ON TOUGHNESS?

Various types of notch toughness and hardness tests havebeen conducted on heat-straightened steel by Rothman(1973), Rothman and Monroe (1973), Pattee et al. (1970)and summarized by Avent (1989). In general, little signifi-cant change was found using Charpy V-notch or RockwellHardness tests. Only the quenched and tempered steelsshowed evidence of small reductions in toughness.

ARE THERE TESTS OR ACCEPTANCE CRITERIAFOR HEAT STRAIGHTENING?

Since hardness tests may be inconclusive, there are cur-rently no nondestructive testing procedures to determine thecondition of the steel after heat straightening. Visualinspection during and after repair remains the most viableoption. It is particularly important to monitor the heatingtemperature, as over-heating is the most likely cause ofdamage to the steel. Other visual signs of possible prob-lems include pitting and blackened spots on the surface.Properly heated steel will have a silvery color after cooling.

DOES MEMBER SHORTENING OCCUR DURINGHEAT STRAIGHTENING?

The subject of member shortening due to heat straighteninghas been mentioned in the literature but little research hasbeen conducted. One researcher stated that using smallervee depth ratios should result in less member shortening,given any particular damage situation, (Moberg, 1979).However, it could be argued that less shortening wouldoccur when using full-depth vee heats, since the top fibershave been heated and are subjected to a tensile stress. Infact, the amount of shortening in a member can be quite sig-nificant, regardless of the vee depth used. If a plate is dam-aged about its strong axis with a midpoint loading, the topedge of the plate experiences compressive yielding (short-ening) and the bottom edge of the plate experiences tensileyielding (stretching). As the plate is subjected to the heatstraightening process, the top edge experiences some“restretching” in the longitudinal direction. However, thesepositive strains are small in comparison to the simultaneousshortening of the bottom edge of the plate. Tests indicatethat shortening at around 0.10 inches occurs during thestraightening of W6×9’s with a degree of damage, ϕd = 6°.Tests on plates have indicated that up to one-half inch ofshortening can occur for ϕd = 25°. Since member shorten-ing may be significant it should be anticipated when con-ducting repairs.

HOW DOES HEAT STRAIGHTENING AFFECTFATIGUE LIFE?

Only one series of fatigue tests on flame straightened mem-bers was found in the literature (AREA, 1946). In this casethree eye bars of A7 steel were heat shortened and thenfatigue cycled. When compared to similar specimens,which had not been heated, the fatigue strength at both500,000 and 1,000,000 cycles was similar. Although data issparse, there is no indication that carbon steels will have ashortened fatigue life after heat straightening.

WHAT TYPE OF RESIDUAL STRESES OCCURAFTER HEAT STRAIGHTENING?

Although residual stresses are often mentioned in literatureon heat straightening, there has been little documentedresearch in this area. Past research was conducted in thecontext of heat curving (not heat straightening), and thus issomewhat limited in its applicability to heat straightening.Some of the most notable research was conducted at theUniversity of Washington (Roeder, 1986), where a finiteelement model was developed to predict the local behaviorof a plate element subjected to a vee heat. Residual stresseswere estimated using the model and experimental strainswere also measured. An example of Roeder’s results are


shown in Figure 14 where tension stresses are positive inthis and other figures of residual stress distribution.

Experimental research was conducted by Brockenbrough(1970b) to back up earlier theoretical residual stress studies,(Brockenbrough, 1970a) on heat-curved plate girders sub-jected to line heats. These stresses, determined by the “sec-tioning method,” were reasonably consistent with thetheoretical values. Similar theoretical methods were usedon vee-heated plate elements by Nicholls and Weerth(1972) and on wide flange beams by Horton (1973). How-ever, the results were not supported by any experimentaldata. Avent et al. (1993a) and Avent and Mukai (1998) con-ducted a study on residual stresses that included both platesand rolled shapes; variations in vee angle, vee depth, andlevel of restraining forces; and degree of initial damage.Residual stress patterns were determined by using the sec-tioning method (Avent and Wells, 1982). The change instrain before and after heating was measured on small stripscut from the heated zone. The strains were converted tostresses using an E = 29,000 ksi.

Shown in Figure 15 are residual stress patterns for ini-tially undamaged plates, which were vee heated four times.A four-inch (100mm)-gauge length centered in the heatedzone was used. For the eight plates tested, relatively littledifference in residual stresses were found for: vee anglesbetween 20 to 82°, moments due to jacking forces betweenzero and 50 percent of member capacity, and vee depthsbetween 3/4 and full depth of the plate. The average valuesare shown in Figure 15 and were similar to those obtainedby Roeder (1985) as shown in Figure 14.

A second series of tests were conducted on plates whichhad been damaged over the range of 6 to 24° correspondingto a strain ratio, µ, ranging between 30 and 100. The num-ber of heating cycles required to straighten the plates rangedfrom 25 to 100. The distribution of residual stresses at the

center of the heated zone was similar to Figure 15. Thelarge number of repetitive heats tended to slightly reducethe residual stresses in comparison to those of the undam-aged plates having only four heats. Overall, maximumcompression residual stresses were approximately 45 per-cent of yield at the flange tips for either case and maximumtensile stresses varied from 20 to 30 percent of yield at thecenter of the plate.

The maximum residual stresses tend to be higher inrolled shapes. Typical residual stress distributions areshown in Figures 16 through 19 for typical rolled shapeswith a negative sign indicating compression. For angles,high compression residual stresses (approaching yield insome cases) were found at the toes and heel as shown inFigure 16. Similarly high values were found at the heelsand toes of channels (Figure 17). Note that strips were notmeasured at the web-flange junctures resulting in disconti-nuities in the graphs. For the Category W heating pattern,wide flange beams tended to have tensile stresses acrossmost of both flanges and compression stresses in the web,as shown in Figure 18. When the beam was re-damagedand heat straightened multiple times, the residual stresseswere mitigated. For Category S damaged wide flangebeams, the pattern reversed with compression stresses inboth flanges and tension in the web, as shown in Figure 19.Residual stresses were measured for a single W6×9 beamwith Category S damage which was repaired using the stan-dard pattern. The maximum compressive stresses in theflanges approached yield while those in the web were some-what less. A comparison of the residual stresses for theundamaged and damaged beams showed a reasonably goodcorrelation for Category S.

The large residual stresses created during heat straighten-ing have several implications. First, if the member is acompression element, the high residual stresses are similar


Fig. 14. Experimental strain and theoretical residual stress distributionfor 2/3 depth, 45° vee heated plate subjected to 1000°F temperature

(Roeder, 1985).Fig. 15. Average residual stress values for vee heated plates which were

originally undamaged.

to welded built-up members. Since U.S. codes use a singlecolumn curve concept, these members are all treated thesame and no capacity reduction would be assumed. How-ever, if multiple column curves are used (typical of manyEuropean countries), then heat straightened columns wouldfall in a lower strength curve after heating due to residualstresses. Consequently, there would be some loss of designstrength.

Second, high tensile residual stresses may reduce theeffectiveness of jacking forces by effectively canceling outthe restraining compression associated with jacking. Move-

ment could be reduced or even reversed, if the jacking forcemoment does not compensate for the residual stresses.

DOES THE DEPTH OF THE VEE INFLUENCE THEAMOUNT OF MOVEMENT DURING HEAT

STRAIGHTENING?

Past researchers (Nicholls and Weerth, 1972; Roeder, 1985)have concluded that the plastic rotation (change in ϕd afterone heating cycle) is proportional to the depth ratio, Rd,which is the ratio of vee depth, dv to plate width, W. Areview of Roeder’s test data in the range of 650°C (±80°) or


Fig. 18. Residual stress distribution damaged, Category W wide flangebeams (assumed E = 200,000 MPa or 29,000 ksi).

Fig. 17. Stresses in channel IX-6 (45° vee, Mj/Mp = 0.50, depth ratio = 1.00).

Fig. 16. Stresses in angle L4×4 (45° vee, apex at heel, Mj/Mp = 0.50, depth ratio = 1.00).

Fig. 19. Residual stresses in Category S damaged wide flange beam (45° vee, Mj/Mp = 0.50, depth ratio = 1.00).

1200°F (±150°) is inconclusive as to vee depth effect. Rec-ognizing that the data was sparse, neither the depth ratio of0.75 nor 0.67 produced plastic rotations that were consis-tently proportional to vee depth. To further evaluate thisbehavior, a series of tests was conducted for depth ratios of0.5, 0.75, and 1.0 and vee angles ranging from 20° to 60°.At least three heats were conducted on initially straightplates for each case and the results averaged. The resultsare shown in Figure 20 for a combination of three depthratios, three vee angles and two jacking ratios. The jackingratios reflect that a jacking force was used to create amoment at the vee heat equal to either 25 percent or 50 per-cent of the ultimate bending capacity of the plate. As canbe seen from Figure 20, the depth ratios of 75 percent and100 percent track each other well. In fact the 75 percentdepth ratio resulted in slightly larger plastic rotations in allbut one of the six cases. The 50 percent depth ratio resultedin erratic behavior when compared to the other two. Inthree of the six cases the 50 percent depth ratio producedmuch smaller plastic rotations. In the other three cases, theplastic rotations were similar.

To further verify this behavior, a series of plates wasdamaged and straightened. The degree of damage was largeenough that at least 20 heats were required for most of theseplates. Therefore, more statistically significant averageplastic rotations were obtained from these tests. Again thepattern of plastic rotations does not have a direct correlationto the vee depth ratios.

Therefore, even though it would seem intuitive thatincreasing the vee depth would increase the plastic rotation,there is no experimental justification for such a generalstatement. It can be concluded that the variation of veedepth ratio between 0.75 and 1.0 has little influence on plas-tic rotation. However, a vee depth ratio of 50 percent mayreduce the plastic rotations.

HOW DOES THE GEOMETRY OF THE PLATEELEMENT EFFECT HEAT STRAIGHTENING?

Researchers have generally considered plate thickness tohave a negligible affect on plastic rotation. The only reser-vation expressed has been that the plate should be thinenough to allow a relatively uniform penetration of the heatthrough the thickness. The practical limiting value is on theorder of 19 to 25 mm (3/4 to 1 in.). Thicker plates can beheated on both sides simultaneously to ensure a uniformdistribution through the thickness or a rosebud tip can beused. Similarly, researchers such as Roeder (1985) haveheated plates of various widths while keeping other param-eters constant. These tests showed no clear relationshipbetween the plastic rotation and plate width.

CAN THE EFFECT OF HEATING TEMPERATUREON DEGREE OF STRAIGHTENING

BE QUANTIFIED?

One of the most important and yet difficult to controlparameters of heat straightening is the through-thicknesstemperature of the heated metal. Factors affecting the tem-perature include: size of torch orifice, intensity of the flame,speed of torch movement, and thickness of the plate. In hisexperiments Roeder (1985) made careful temperaturemeasurements of the heats produced by knowledgeablepractitioners. He found that these individuals, when judg-ing temperature by color, commonly misjudged by 56°C(100°F) and, in some cases, as much as 111°C (200°F).Thus, there are considerable variations in temperature con-trol, even with knowledgeable users.

Assuming adequate control of the applied temperature ismaintained, the question arises as to what temperature pro-duces the best results in heat straightening without alteringthe material properties. Previous investigators have dif-fered in answering this question. For example, Shanafeltand Horn (1984) state that heats above 650°C (1200°F) oncarbon and low alloy steels will not increase plastic rota-tion. Rothman and Monroe (1973) concluded that reheat-ing areas where previous spot heats were performed wouldnot produce any useful movements. However, Roeder(1985) has shown that the resulting plastic rotation is gen-erally proportional to the heating temperature up to at least870°C (1600°F). To more clearly define the behavior sug-gested by a limited number of data points in Roeder’s study,a series of heats were applied to plates in which the heatingtemperature was varied from 370 to 815°C (700° to1500°F) in increments of 56°C (100°F). The results asshown in Figure 21 establish a clear and regular progressionof increased plastic rotation with increasing temperature.Because increasing the temperature leads to more move-ment, practitioners sometimes tend to overheat. As dis-cussed previously, it is important to limit the maximum


0

1

2

3

4

5

6

7

8

9

0.5 0.75 1 0.5 0.75 1 0.5 0.75 1 0.5 0.75 1 0.5 0.75 1 0.5 0.75 1

Pla

stic

Ro

tati

on

(m

illir

adia

ns)

Depth Ratio

Θ 20 45 60 20 45 600 0 0 0 0 0

Jacking Ratio 0.5 0.5 0.5 0.25 0.25 0.25

Fig. 20. Influence of vee depth on plastic rotations of originally straightplates for various vee angles and jacking ratios

(heating temperature = 650°C or 1200°F).

temperature in order to insure that the material propertiesremain unchanged.

HOW DOES THE LEVEL OF JACKING FORCEINFLUENCE THE DEGREE OF STRAIGHTENING

PER HEAT?

While practitioners have long recognized the importance ofapplying jacking forces during the heat-straighteningprocess, little research has been conducted to quantify itseffect. A series of tests designed to evaluate this parameterinvolved applying a jacking force to a plate such that amoment is created about the strong axis in a direction tend-ing to close the vee. This moment (at ambient temperature)is non-dimensionalized for comparison purposes by form-ing a ratio of the moment at the vee due to the jacking force,Mj, to the plastic moment, Mp, of the cross section, that isMj/Mp. This term is referred to as the jacking ratio. Thetests included jacking ratios ranging from zero to 50 percentwith four different vee angles and the vees extending overeither 3/4 or full depth of the plate. The results shown in Fig-ure 22 indicate how jacking can increase movement percycle.

Roeder (1985) also studied the effect of the jacking ratiovariation and found a similar pattern of behavior. However,the number of data points was limited. In general, theamount of movement is proportional to the jacking ratio.Experimental evidence indicates that this relationship istrue for various cross sectional shapes. Several examplesare shown in Figures 23 to 25. This data shows that therestraint necessary during heating can result from: jackingforces, the temperature gradient of the cool surroundingmaterial, or a combination of both. For any combinationproducing the same restraint, the effect on the steel is simi-lar. For example, in vee heated elements, advantage is

taken of both material and jacking constraints. For lineheats in plates (particularly thin plates), jacking dominatesbecause thermal gradients are small.

CAN DAMAGED STEEL BE REPAIRED MORETHAN ONCE BY HEAT STRAIGHTENING?

A series of Category W damaged beams were re-damagedand straightened up to eight times, Avent, et. al., (1993a).The beams had a degree of damage of approximately 7°.While the beams were successfully straightened, materialproperties were noticeably affected by the fourth cycle ofdamage and repair. Both yield and tensile strengthincreased, especially in the vicinity of the apex of the veeheats. The difference between tensile strength and yielddecreased with each damage-repair cycle indicating a trendtoward more brittleness. Ductility also decreased signifi-cantly with successive damage-repair cycles. In anotherseries of tests on composite girders (Avent, Madan, andShenoy, 1993b), a crack formed along a line heat during thefourth damage-repair cycle.

Based on these results it is recommended that the sameregions not be heat straightened more than twice if re-dam-aged. Should additional damage occur, the changes inmechanical properties as well as the possibility of crackingshould be taken into account when deciding whether heatstraightening is more viable than member replacement.

CAN THE AMOUNT OF STRAIGHTENING PERHEAT BE PREDICTED?

Two general approaches have been used to develop an ana-lytical procedure for predicting member response during aheat-straightening repair. One approach involves finite ele-ment/finite strip thermal and stress analyses including


Fig. 21. Influence of heating temperature on plastic rotation for 3/4 depthvee heats and a jacking ratio of 0.16.

0

1

2

3

4

5

6

7

8

9

10

11

12

13

14

0 10 20 30 40 50 60 70 80 90

Vee Angle (degrees)

Pla

stic

Ro

tati

on

(m

illir

adia

ns)

Fig. 22. Influence of jacking ratio on average plastic rotation for ¾depth vee heats and 650°C (1200°F) heating temperatures

(lines represent a least squares curve fit).

inelastic behavior (Roeder, 1985). The stress and strainequilibrium is evaluated over small time steps and takesinto account the influence of the non-uniform temperaturedistribution. This approach is a lengthy computational task,which is only possible using computer techniques and a typ-ical analysis for a single vee heat can require extensive setup and computer time.

The other approach considers the global action of thevee. The Holt equation is based on such an approach andassumes that perfect confinement is provided at all timesduring the heating phase. As a result, longitudinal dis-placements through the vee are linear. With this equationthe number of vee heats required to remove a bend in a steelmember can be simply calculated. However, the Holt equa-tion neglects the effect of restraining forces and temperaturevariation. A set of assumptions that are consistent with theobserved behavior of vee heated flexural members include:(1) longitudinal plastic strain occurs only in the middle two-thirds of the vee heat zone (and in a reflected vee about theapex for partial depth vees); (2) at any specified distancefrom the neutral axis of the plate, the strains are constant inthe longitudinal direction over the middle two-thirds of thevee; (3) the planes defined by the sides of the vee remainplane after heating and rotate about the apex of the vee; (4)the plastic rotation per heat varies linearly with the jackingratio, Mj/Mp; (5) the degree of internal confinement (exclud-ing any jacking force) equals 60 percent of perfect confine-ment; (6) the maximum heating temperature averages1200°F (650°C); (7) the plastic rotation per vee heat varieslinearly with the ratio of the product of the stiffening ele-ment width, bs, and its distance from the apex of the vee, ds,to the square of the vee depth, d; and (8) the movement pervee heat is a linear function of the ratio of plastic to elasticsection modulus, Z/S, about the axis of bending. Usingthese assumptions, an expression for the plastic rotation, ϕ,per vee heat for Category S or W damaged wide flangebeams or channels or angles bent about their x or y axis isgiven by

where , the factor associated with the external jackingforce is given by

Fs, a factor reflecting the shape of the cross section, isgiven by

Fa, a stress factor, is given by


0

1

2

3

4

5

6

0 10 20 30 40 50

Vee Angle (degrees)

Pla

stic

Ro

tati

on

(m

illir

adia

ns) Jacking Ratio=0.0

Jacking Ratio=0.25

Jacking Ratio=0.50

Theoretical Eq. 6.1

Least squares curve fit

Jacking Ratio=0.50

Jacking Ratio=0.25

Jacking Ratio=0.0

Fig. 23. Experimental and theoretical plastic rotations for a C 6x8 channel with Category S damage.

0

5

10

15

20

25

0 10 20 30 40 50

Vee Angle (degrees)

Pla

stic

Ro

tati

on

(m

illir

adia

ns) Jacking Ratio=0.0

Jacking Ratio=0.25

Jacking Ratio=0.50

Theoretical Eq. 6.1

Least squares curve fit

Jacking Ratio=0.50

Jacking Ratio=0.0

Jacking Ratio=0.25

Fig. 24. Experimental and theoretical plastic rotations for Category Wdamage of a C 6x8.2 with the open end of vees at flange-web-juncture.

0

1

2

3

4

5

6

7

8

9

10

0 10 20 30 40 50

Vee Angle (degrees)

Pla

stic

Ro

tati

on

(m

illir

adia

ns)

Jacking Ratio=0.0

Jacking Ratio=0.25

Jacking Ratio=0.50

Theoretical Eq. 6.1

Experimental least squares curve fit

Jacking Ratio=0.50

Jacking Ratio=0.25

Jacking Ratio=0.0

Fig. 25. Plastic rotation versus vee angle for W6x9 using the CategoryW heating pattern (Temperature = 650°C or 1,200°F).

ϕ ϕp s a bF F F= l (7)

FM

Mj

pl= +0 6 2. (8)

Fb d

ds

s s= +11

2 2( ) (9)

FZ

S

M

Maj

p

= − − FHGIKJFHGIKJ

LNM

OQP

1 2 12

3(10)

Fl

and ϕb is the basic plastic rotation factor for a rectangularplate having no external jacking force and is given by

HOW DO AXIALLY LOADED COMPRESSIONMEMBERS RESPOND TO HEAT

STRAIGHTENING?

For columns and axially loaded members it is important toconsider the P-∆ effect. If an axially compressed memberis damaged by lateral loads as shown in Figure 26, amoment is generated which is equal to P∆. This moment isin the opposite direction to the moment generated by a jack-ing force during the straightening process. The moment dueto the P∆ will oppose the restoration movement during heatstraightening.

A series of full-scale tests were conducted on both Cate-gory W and S damaged axially loaded columns (Avent andMukai, 1998). Seven straightening procedures were con-ducted with at least 10 heats used for each set of parame-ters. The axial load imposed during the straighteningprocess was typically 35 percent of the design allowablevalue. The jacking force was modified to cancel the P∆moment effect in addition to providing the specified jackingratio. The response for both cases was significantly lessthan that predicted for a wide flange beam without axialload. From these tests it is concluded that heat straighten-ing is effective for axially loaded members using the sameheating patterns as for cases without axial loads. However,the movement per heat will be somewhat smaller than if noaxial loads are present.

WHAT HEATING PATTERN SHOULD BE USEDFOR LATERAL IMPACT OF THE LOWER

FLANGE OF A COMPOSITE GIRDER?

Perhaps the most typical type of damage found on steelbridge members results from impact of vehicles or freight

on the beams or girders of composite deck-girder bridges.Heat straightening is an attractive repair alternative becauseof its low cost and minimal disruption of traffic (Avent andBrakke, 1996). Typical damage includes: (1) a flexuralyield zone at the impact point in the lower flange due tobending about the strong axis of the flange; (2) a yield lineor yield zone in the web due to bending of the web elementabout its weak axis; (3) reverse curvature flexural yieldzones in the lower flange at diaphragms on either side of theimpact point; and (4) localized bulges in the lower flangeand web. Conceptually, vee heats are used to repair plateelements with plastic bending about the major axis, whileline heats are applied to repair plate elements with flexuraldamage about the minor axis. Hence, a vee heat on the bot-tom flange in conjunction with a line heat on the web,applied to their respective plastically yielded portions, arethe proper heat patterns to repair composite beamsimpacted by high loads (Figure 27). Care must be taken tocontinually adjust the span of the line heats, so that onlythose portions of the web are heated that show plastic cur-vature after the last heating cycle. Similarly, the vee heatsare confined to the portion of the bottom flange with plasticdeformations. In addition, a half-depth web strip heat isapplied. The purpose of this heat is to reduce the differen-tial shortening between web and flange. By heating theweb with a half-depth strip, the web can deform and relievesome of these stresses. The application of strip heats on theweb do not influence the average plastic rotations apprecia-bly but, does tend to reduce the buckling of the web near thecenter of damage.

HOW CAN THE LIMIT FOR SAFE JACKINGFORCES BE DETERMINED FOR DAMAGED

COMPOSITE BEAMS?

When a jacking force is applied to the lower flange of acomposite beam, the internal redundancy makes it difficultto calculate the stress in the lower flange. Such difficultiesmay occur on other types of indeterminate structures as


ϕ θb = 0 0147

3. sin (11)

Fig. 26. P∆ effect on an axially loaded column. Fig. 27. Heating patterns for composite girder.

well. The question is: how much of the jacking force istranslated into lower flange moment and how much is trans-ferred through the web and into the deck? When applyinga lateral load to the bottom flange near the center of thecomposite girder, the moment produced is transferred to theend reactions by two mechanisms: the bottom flange acts asa flexural beam supported at the ends, and the web acts as aflexural plate (and, at large deformations, a membraneplate) supported by the deck and end diaphragms. Asdemonstrated by Avent and Fadous (1989), a shallowergirder has greater stiffness and thus lower flange momentsthan that of the deeper girder subjected to the same lateralload. By adopting a jacking ratio definition which assumesthat all of the applied force is transferred into a lower flangemoment, Mj, which is resisted only by the plastic momentof the bottom flange, Mp, the implication is that most of thelateral stiffness is provided by the bottom flange. Thisapparent jacking ratio can be misleading (Avent et al.,1993b). It is more relevant to use an effective jacking ratio,Mf, which reflects the actual moment in the flange to itsplastic capacity, Mp. This relationship has been developedas a function of the ratio of beam depth, d, to web thicknesstw in the form

where

Based on sudden fractures that have occurred at higherjacking ratios, it is recommended that the jacking ratioMf/Mp be limited to one-third.

CAN THE MOVEMENT PER HEAT OFCOMPOSITE BEAMS BE PREDICTED?

An equation for plastic rotation has been developed byAvent and Mukai (1998) in the form of that for rolledshapes, Equation 7, where ϕb is defined by Equation 11, Fs = 1.0, γ by Equation 13, and

and

or

HOW CAN LOCALIZED DAMAGE BE QUANTIFIED?

The focus of past heat straightening research has been onvarious aspects of repairing global damage. However, it isa rare situation when localized damage doesn’t occur con-currently with global damage. Yet, little published infor-mation has been available on how to repair local damage byheat straightening. As a result, localized damage is oftenrepaired improperly by cold mechanical straightening andhot mechanical straightening, as well as heat straightening.

Local damage patterns display two main characteristics:large plastic strains (usually tensile) in the damaged zone,and bending of plate elements about their weak axis. If thelocal damage is to be repaired, shortening must be inducedin the damaged area equal to the elongation caused whenthe element was damaged. In addition, the distortion alongthe yield lines must be removed as part of the repairprocess. Studies on global damage repair have shown thatvee heated regions shorten significantly during cooling andthat line heats can be used to induce bending about the yieldlines. Thus a combination of line and vee heats can be usedto repair localized damage. It is helpful to separate local-ized damage into two distinct categories: Category L/Sdenoting bulges in plate elements with stiffening elementson two sides (such as a web), and Category L/U denotingbulges in plate elements having one unstiffened edge (suchas a flange). Examples are shown in Figure 28.

HOW SHOULD LOCALIZED DAMAGE TOUNSTIFFENED ELEMENTS BE

HEAT STRAIGHTENED?

While every damaged unstiffened plate element will haveits own unique shape, there is a set of common characteris-tics, which can be used to determine the heating pattern.


Fig. 28. Typical localized damage classified as Category L.

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15 2 75,

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Fd t

aw= FHGIKJ

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(14)

FM

Mf

pl = +FHG

IKJ

0 6 2.

FM

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(16)

Category L/U damage is commonly caused by impact toa flange on one side of the web. An idealized zone of dam-age is shown in Figure 29. The impacted side of the dam-aged flange will be referred to as the near side (N). Thenon-impacted side of the flange on the other side of the webwill also typically incur damage. This damage on the farside of the flange (F) has geometry similar to that of type Nbut usually of lower magnitude. The damaged flange typi-cally undergoes rotation about a clearly defined yield-linenear the fillet of the web. The impacted side of the flange(side N) usually deforms in a folded plate pattern. Thisflange is shown deforming toward the web in Figure 29b.The deformation usually results in yield lines, which definethe edges of the folded plate (Figure 29c). In some cases,particularly in regions of high curvature, the deformationpattern may be one of a flexural yield zone rather than aseries of yield lines. These zones result from plate elementflexure and tend to spread over the surface as the degree ofdamage increases. Such zones will be referred to here asyield surfaces. The other half of the same flange usuallydeforms in a similar pattern in the opposite direction, evenif not directly impacted. The folded plate pattern, Figure29d, tends to have smaller deformations, thus δn > δf.Because the web is thinner than the flange, a yield line oftenforms in the web near the fillet. The section shown in Fig-ure 29b illustrates this behavior. The tee section at theflange/web juncture remains a rigid right angle. The yieldline forming in the web fillet allows this tee to rotatethrough an angle θw. The yield line at the flange fillet on theimpacted side of the flange (side N) results from the addi-tional rotation, θn, thus the total rotation of the N flange isθw + θn. The other half of the flange (side F) tends to act asa reaction against rotation thus a second flange yield line

may form at the F side fillet. The angle formed by this yieldline is θf and the rotation of the F flange is θw - θf. The iden-tification of these yield lines is important in the repair pro-cedure.

The heating/jacking pattern to straighten this damagewill depend on how the geometry changes as heat straight-ening progresses. It is usually most effective to have jack-ing forces on both the near and far sides of the flange.However, straightening can be conducted with jacking onlyon the near (impacted) side. The specific steps are:

1. Restraining forces

Place jacking forces on both the near and far sides of thedamaged flange in the direction tending to restore theflange to its original condition. As shown in Figure 30a,a convenient arrangement on the near side is to place ajack, Pn, between the top and bottom flange. The far sidejack, Pf, requires a clamping type force which is oftenmore difficult to arrange in field applications. If theclamping force cannot be anchored from the oppositeflange, a spreader beam arrangement can be used, asshown in Figure 30d, to anchor the reaction to thestraight portions of the far side flange. An alternative isto only jack from the near side.

However, the average movement per cycle tends to belower than similar cases jacked on both sides. Careshould be taken to limit the magnitude of jacking forcesused.

2. Vee heats

Although vee heats may not be necessary, a limited num-

ENGINEERING JOURNAL / FIRST QUARTER / 2001 45

Fig. 29. Heat straightening local flange damage (Category L/U).

ber may be used to assist in the flange shortening effort.The vees should be approximately half depth and appliedto both the near and far sides of the flange to eliminateglobal curving of the member. The vee should be narrowwith an angle of 20° or less and the open end of the veesshould be at the flange tips. It is best to place the veeheats in regions where no line heats are required. Nomore than two vees should be used (preferably only one)in one heating cycle. The location should be shifted witheach heating cycle so the same location is not re-heatedfor at least three cycles. A typical arrangement is shownin Figure 31b.

3. Line heats

A pattern of line heats should be placed so that the lineheats correspond to yield lines. These should be heatedon the convex surface if possible. A typical pattern isshown in Figure 31a. For yield surfaces of continuousplastic strain, such as often occurs in regions such asABC in Figure 31a, line heats should be spaced over the

section at a spacing of approximately bf/4 where bf is theflange width. Similarly, line heats may also be usedinstead of vee heats on section BCDE. The order of heat-ing the yield lines tends to have a minor impact althoughit is good practice to heat the ones at the largest damagelocations first. It is also recommended to heat the nearside lines prior to the far side.

4. Web line heat

The web yield line should be heated last. It is typicallylocated at the fillet as shown in Figure 31c.

These four steps complete the cycle. The cycle shouldbe repeated until the flange is straight.

Quite often this procedure can be used to nearlystraighten the section. However, the progress of themovement should be observed to insure that over-straightening does not take place. If the near flangemovement progresses too quickly, then θn may becomezero prior to θw. This situation is shown in Figure 30b.Should this behavior occur, a modification should bemade in Step 3 for line heats. Rather than heating alllines on the near side (31a), line 4 should not be heated.

If straightening progresses to the point that θf = θw,then the far flange may over-straighten with the continu-ation of heating. The pattern should be changed. The sit-uation is depicted in Figure 30c. The modification is toreverse the direction of the far side jacking force whilecontinuing the pattern. The force Pf will prevent over-


Fig. 30. Arrangement of restraining forces during various stages of repair. Fig. 31. Arrangement of vee and line heats.

straightening while allowing the near flange and web tocontinue corrective movement. Judgement is required inselecting the number of line heats to fit the specific dam-age. As straightening progresses, the number of linesmay be adjusted to fit the changing shape of the damage.

HOW SHOULD LOCALIZED DAMAGE TO STIFFENED ELEMENTS BE

HEAT STRAIGHTENED?

Stiffened elements of a beam are defined as those supportedon both sides by perpendicular elements. The web of awide flange beam is a typical example in that the twoflanges provide the support or stiffening effect. When stiff-ened elements are damaged, the pattern formed is usually adish shaped bulge. This damage frequently approximates ashallow spherical shell.

Judgement is required when selecting the heating pat-terns for damaged stiffened elements. Heating the regionsof sharpest curvature with combinations of lines and/or nar-row vees is the most effective approach. Always follow thepractice of heating only in the vicinity of regions with plas-tic curvature. As straightening progresses, this region maybecome smaller. The following methodology is recom-mended for bulges in stiffened elements. The jacking forceis typically applied at the crown of the bulge using aspreader plate.

1. Initial Heating Pattern

The typical bulge will have reverse curvature bending asshown in Figure 32. The crown region should be heatedfirst with the torch on the convex side. As movementprogresses, the heating patterns can be expanded into thereverse curvature region again with the torch on the con-vex side. The initial heating patterns should consist ofradial and ring line heats as illustrated in Figure 32. Theexact number of ring heats will depend on the size of thisregion. It is recommended that the diameter of the small-est ring be no less 51 mm (2 in.) and that spacing betweenrings be at least 51 mm (2 in.). For large bulges the ringspacing should be larger than 51 mm (2 in.). For caseswhere the curvature is relatively uniform, equally spacedrings may be used. However, a ring heat should beplaced at locations where sharp changes in curvature areindicated. Heat the outer ring on the convex side firstand work inward. After the rings are heated, the radiallines should be heated. Again, work from the outside inbut do not run the radial lines inside the last ring. Con-tinue this pattern cyclically until the crown region beginsto flatten. Allow the steel to completely cool betweenheating cycles.

2. Final Heating Pattern

As the crown section flattens, the heating pattern shouldbe expanded into the reverse curvature regions. Onealternative is to expand the number of ring heats andextend the radial heats as described in Step 1 and shownby dashed lines in Figure 32b. The second alternative isto use four-point star vee heats along with ring heatsinstead of radials. The star vee pattern is shown in Fig-ure 33. The vees are heated first, working from tip tocenter. However, the central portion past where adjacent


Fig. 32. Curvature and line heating patterns for category L/S damage.

Fig. 33. Star vee heat pattern.

vees intersect should not be heated. Vees should be nar-row with an angle of 20° or less. After star vee heating,the rings should be heated, starting with the outermostring. A maximum of three ring heats should be used atone time with the star vees. Rotate the star vee pattern30° after each cycle so that the same vees are not repeti-tively heated. Rings may be repetitively heated orshifted, depending on the degree of plastic curvature.The steel should completely cool before the next heatingcycle begins.

WHAT TYPE OF TOLERANCES SHOULD BEUSED IN HEAT STRAIGHTENING?

In most cases any desired tolerance level can be obtained byheat straightening. One approach to specifying the toler-ance for heat straightening repair is to use a standard spec-ification for manufacturing or fabrication. One commonlyreferenced standard is the American Welding SocietyBridge Welding Code (1996 or latest edition). Memberdimensional tolerances are given in Section 3.5 and flatplate tolerances are given in Section 9.19. The use of thisor similar construction standards for heat straightening tol-erances may be too limiting for the following reasons: (1)The tighter the tolerances, the longer the repair will take.Costs and delays in using the structure will thus increase.(2) In many cases, the area of damage is small and not nec-essarily located at the most highly stressed region. Lessrestrictive tolerances can safely be used in such instancessince the engineer can assess the specific structural config-uration. (3) The steel in the vicinity of impact may havematerial distortions, such as changes in thickness, whichmay preclude highly restrictive tolerances. As a conse-quence, engineering judgement for the specific situation isrecommended as the most effective approach in setting tol-erances.

In order to provide some general guidelines, a set of rec-ommended tolerances is given in Table 1. These tolerancelevels are less restrictive than specified in the Bridge Weld-ing Code.

SUMMARY AND CONCLUSIONS

The goal of this paper has been to summarize the state-of-the-art in heat straightening repair of damaged steel. Theformat used was to list the commonly asked questions andto provide the answers based on the latest research anddevelopment. The principles discussed here can also beapplied to heat curving members to produce camber, sweepor curved beams. Heat straightening is a safe and econom-ical repair procedure when properly executed. It is a skillrequiring practice and experience. The proper placementand sequencing of heats combined with control of the heat-ing temperature and jacking forces distinguishes the expertpractitioner. In the past, heat straightening has been moreart than science. With the information presented here, thepracticing engineer will have guidelines quantifying themost suitable procedures.

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American Welding Society (1996), Bridge Welding Code,ANSI/AASHTO/AWS D1.5-96, Miami, FL.

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Avent, R. R. (1995), “Engineered Heat Straightening,” Proceed-ings, National Steel Construction Conference, AISC, San Anto-nio, TX, May.


Table 1. Recommended Tolerances for Heat Straightening Repair

Member Type Recommended Minimum Tolerance1,2

English (in.) SI (mm)

Beams, Truss Members, or Columns: overall sections including

impact point

½ in over 20 ft ¾ in over 20 ft

13 mm over 6 m 19 mm over 6 m

Local Web Deviations d/100 but not less than ¼ in. d/100 but not less than 6 mm

Local Flange Deviations b/100 but not less than ¼ in. b/100 but not less than 6 mm 1Units of member depth, d, and flange width, b, are inches and millimeters, respectively, for English and SI units

2Tolerances for curved or cambered members should account for the original shape of the member

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Ciesicki, H., and Butler, S. (1968), “Flame and Camber Your OwnBeams,” Industry and Welding, 31(4), pp. 76-79.

Welding Engineer (1959), “Flame Buckled This Steel and FlameStraightened It, Part 1,” Welding Engineer, 44(2), p. 43.

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Horton, D. L. (1973), “Heat Curved Mild Steel Wide Flange Sec-tions: An Experimental and Theoretical Analysis,” Thesis pre-sented to the University of Washington, at Seattle, WA, inpartial fulfillment of the requirements for the Master of Sciencedegree.

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