Evaluating Fire-Damaged 

Components of Historic 

Covered Bridges

Brian Kukay Logan Dunlap

Charles Todd Robert White

Tyler Jahn Mark Dietenberger

Jenson Sannon

United States Department of Agriculture

Forest Service

Forest Products Laboratory

General Technical Report

FPL–GTR–243

August 2016

In cooperation with the 

United States Department of Transportation

Federal Highway Administration

A FHWA-NHCBP (National Historic Covered 

Bridge Program) publication
Abstract

Arson continues to claim many historic covered bridges. 

Site-specific, post-fire evaluations of the structural integrity of a bridge are often necessary in a fire’s aftermath. 

Decisions on whether individual wood components can be 

rehabilitated, reconstructed, or replaced must be made. This 

report includes coverage of existing approaches and exploratory approaches that can be used for general guidance 

and a select number of more specific treatments that can be 

used to approximate the residual capacity of individual firedamaged members. Topics such as fire-retardant treatments 

and other measures to prevent future fire damage are also 

discussed.

Keywords: covered bridge damage, rehabilitation of fire 

damaged bridge, fire resistant protection of structural wood

August 2016

Kukay, Brian; Todd, Charles; Jahn, Tyler; Sannon, Jenson; Dunlap, Logan; 

White, Robert; Dietenberger, Mark. 2016. Evaluating fire-damaged components of historic covered bridges. General Technical Report FPL-GTR-243. 

Madison, WI: U.S. Department of Agriculture, Forest Service, Forest Products Laboratory. 34 p.

A limited number of free copies of this publication are available to the 

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Acknowledgments

The authors acknowledge the many contributions of James 

Wacker, Sheila Rimal Duwadi, Mike Ritter, Carol Clausen, 

and Bob Ross.

Cover photo: The Cedar Covered Bridge, Winterset, Iowa, 

was destroyed by a fire set by an unknown arsonist on the 

evening of September 3, 2002. The Madison County Board 

of Supervisors have constructed a replica of the original 

bridge. Image courtesy of the Winterset Madisonian/Dave 

Braga.

This study is part of the Research, Technology and 

Education portion of the National Historic Covered 

Bridge Preservation (NHCBP) Program administered 

by the Federal Highway Administration. The NHCBP 

program includes preservation, rehabilitation and 

restoration of covered bridges that are listed or are 

eligible for listing on the National Register of Historic 

Places; research for better means of restoring, and protecting these bridges; development of educational aids; 

and technology transfer to disseminate information on 

covered bridges in order to preserve the Nation’s 

cultural heritage.

This study is conducted under a joint agreement between the Federal Highway Administration–Turner 

Fairbank Highway Research Center, and the Forest 

Service–Forest Products Labortatory.

Federal Highway Administration Program Manager—

Sheila Rimal Duwadi, P.E.

Forest Products Laboratory Program Manager—

Michael A. Ritter, P.E.
Contents

Executive Summary.........................................................................1

1.0 Introduction................................................................................1

2.0 Background Information............................................................2

3.0 Reduced Section Analysis of Fire-Damaged 

Wood Elements................................................................................6

4.0 Post-Fire Inspection and Site Visit.............................................7

5.0 Options for Evaluating Heat-Damaged Wood ...........................9

6.0 Other Post-Fire Assessments......................................................16

7.0 Post-Fire Bridge Repair .............................................................17

8.0 Design for Fire Prevention and Control.....................................19

9.0 Concluding Remarks and Summary ..........................................23

10.0 Federal Program for Covered Bridges.....................................23

References........................................................................................24

Additional Sources...........................................................................29

Appendix—Development of a New NDE Test for 

Fire-Damaged Wood ........................................................................30
In memory of

Robert White,

who will live on through the 

historic covered wood bridges 

that dot our great country, 

through the county bridge 

engineers that make use of this 

report, and through those who 

reference his many other works 

now and in the future.
Evaluating Fire-Damaged Components of 

Historic Covered Bridges

Brian Kukay, Ph.D., P.E, Associate Professor of General Engineering

Charles Todd, Ph.D., Tenured Full Professor, Department of Mathematical Sciences

Tyler Jahn, M.S., Graduate Student, General Engineering 

Jenson Sannon, M.S., Graduate Student, General Engineering

Logan Dunlap, B.S., Student, General Engineering

Montana Tech of the University of Montana, Butte, Montana

Robert White, Ph.D., Research Forest Products Technologist

Mark Dietenberger, Ph.D., Research General Engineer

USDA Forest Service, Forest Products Laboratory, Madison, Wisconsin

Executive Summary

Covered bridges, which used to cover the American landscape, now number less than 900 bridges as a result of various factors including neglect, arson, vandalism, and natural 

disasters. Covered bridges tend to be in isolated locations 

and constructed of flammable materials; therefore, they are 

susceptible to arson and vandalism. Because it is typically 

difficult for firefighters to arrive at the bridge in a timely 

manner, any fire developing on the bridge can cause critical damage or completely destroy the bridge. Therefore, 

numerous security measures have been increasingly applied 

that provide deterrence, deny access, detect threats, provide 

protection, and educate the public. Passive protections such 

as lighting and fire retardant coatings are relatively low 

cost. Additional fire protection technologies are discussed 

in the Guide for In-Place Treatment of Wood in Historic 

Covered and Modern Bridges (Lebow and others 2012b). 

However, arsonists can overcome such protection and will 

require active security measures as described in the Covered 

Bridge Security Manual (Phares and others 2013). Although 

the famous Cedar Bridge in Madison County, Iowa, was 

destroyed in 2002 (cover page), along with many others 

at the same time and earlier, many more covered bridges 

are now surviving fires because of increased security measures implemented over the years. Because of the low costs 

involved, restoration of such limited damaged bridges is 

becoming increasingly viable rather than wholesale replacements, thereby extending the scarce restoration funding for 

the historic structures. Considerable confusion remains for 

understanding the level of fire damage and of the restoration process required. Much of this is not discussed in other 

manuals and is addressed here.

This guide will identify technologies and methodologies 

used to evaluate the fire-damaged components and the consequential methods of restoration and rehabilitation will 

also be identified and organized. The guide is provided in 

the logical order, as first understanding fire damage effects 

on the bridge structural properties (Sections 2 and 3), evaluating the fire-damaged wood elements (Sections 4 to 6), 

restoring the bridge with repairs, fire prevention, and controls (Sections 7 and 8), and citing standards, programs, and 

guidance (Sections 9 and 10). Perhaps the most important 

chapter to engineers is Section 3 on reduced section analysis 

of fire-damaged wood elements that provide a basis to guide 

the later chapters on inspection and restoration. The tools for 

post-fire inspection of the char regions mainly determine the 

depth of char, preferably using a minimal destructive device 

or NDE testing (see Appendix) of the char layer thickness 

determination. By having a value for the char layer thickness as input to the loading analysis of the reduced section 

will help determine if the char layer only needs to be removed or left as it is for posterity, or if the strength and load 

capacity of the damaged member is calculated to be so reduced by a deep char that it needs to be replaced. Although 

the presence of char on the covered bridge can look alarming, a careful post-fire inspection can reveal how minimal a 

restoration is actually required. This would be particularly 

true if the security measures installed provide timely fire 

protection that prevent or suppresses a historic covered 

bridge fire in its early stages.

1.0 Introduction

In 1804, the first documented covered bridge was built in 

America: the Union Bridge by Theodore Burr. This bridge 

was constructed in New York and spanned the Hudson River 

for 105 years (Griggs 2014). Although this bridge no longer 

exists, historic covered timber bridges grace landscapes 

across the country. Neglect, improper maintenance, and fire 

damage from both people and nature challenge the life of 

these bridges. It is important to properly maintain and preserve these historic, sometimes even romantic, landmarks 

as they connect us to the past. A 2012 photograph of the 

historic Red Covered Bridge, completed in 1864, is shown 

in Figure 1.
General Technical Report FPL–GTR–243

2

Covered bridges began to spread westward and soon became a staple of historic transportation. Today, approximately 500 to 600 historic covered bridges are still standing 

in the United States (Wacker and Duwadi 2008). A proactive 

movement has begun to ensure proper rehabilitation, restoration, and preservation of these historic landmarks while 

simultaneously raising public awareness. 

After a historic covered bridge has been affected by fire, a 

post-fire assessment is advantageous to investigate altered, 

individual members and their connections. The vast majority 

of historic covered bridges are made of timber stringers cut 

from solid sawn lumber. Unless the fire is severe, it is usually not necessary to replace all the large members after a 

fire. Selwyn Fox (1974) conducted a study using three 50-ft 

24F Douglas-fir stress-grade beams that were salvaged from 

a warehouse after a severe fire had reduced each member’s 

cross section by nearly 10%. Two-point loading tests were 

conducted on each beam after they were cut into six 10-ft 

sections and three 30-ft sections. The results showed that 

there was no evidence that the fire had reduced the strength 

of the beams beyond the reduced cross section and effected 

glue-line.

The goal for historic covered bridges is to return them to 

their original state while staying true to historic building 

practices and materials. Rehabilitation and reinforcement 

of fire-damaged timber stringers provide a method that is 

much less costly than replacing those damaged parts. This 

is important because bridge replacement may not be an 

option for various reasons (Wacker and Duwadi 2008). If 

appropriate analysis, repairs, and treatment are undertaken, 

fire-damaged wood members can often be restored instead 

of replaced (White and Ross 2014). 

The objective of this publication is to provide guidance on 

the available methodologies to evaluate the residual load 

capacities of fire-damaged wood members. After a review 

of some background information, we discuss the reduced 

section approach normally used to evaluate the residual load 

capacity of a fire-damaged wood member. Next, we discuss 

items that should be included in a site visit of the damaged 

bridge. For situations where uncharred wood needs to be 

evaluated for reduced load capacity, we review the options 

for evaluating the residual wood for reduced mechanical 

properties. We conclude the publication with discussion of 

post-fire repair and actions that can be taken for fire prevention and damage control. 

2.0 Background Information

When heated to high temperatures, wood undergoes thermal 

degradation to char and volatile gases. Ignition of the volatile gases results in the flames that facilitate the spread of 

the fire. Surface heating of a wood member in a fire results 

in surface charring and a steep temperature gradient. The 

stages of thermal wood degradation become zones of degradation in a structural wood member exposed to fire. In a 

Figure 1—Red Covered Bridge near Princeton, Illinois.
Evaluating Fire-Damaged Components of Historic Covered Bridges

3

broad sense, there is an outer char layer, a pyrolysis zone, a 

zone of elevated temperatures, and the cool interior (Fig. 2). 

These zones of degradation reflect the temperature profile 

through the cross section. The extent of thermal degradation 

within these zones determines the residual load capacity of 

the wood components after a fire.

2.1 Thermal Degradation, Ignition, and 

Charring of Wood

Thermal degradation without apparent charring or mass 

losses can be detected first as losses in flexural properties at 

elevated temperatures (White and Dietenberger 2001, 2010). 

From 200 to 300 °C (392 to 572 °F), the wood components 

of extractives, hemicelluloses, and lignin begin to undergo 

significant degradation resulting in mass losses and complete loss of strength (Rowell and Dietenberger 2013; Wang 

and others 2014). Significant depolymerization of the cellulose component of wood occurs between 300 and 350 °C 

(572 and 662 °F). Thermal degradation of the wood to a 

char residue causes reductions in the density of the wood 

and the shrinkage of the surface char layer. The thermal degradation of a wood member depends upon the temperature 

and duration of the fire, with the char layer increasing as 

exposure increases.

Wood can ignite under piloted ignition or autoignition. Autoignition is where wood ignites in the absence of a direct 

flame. Piloted ignition is where an external flame or spark 

can ignite the combustible gases generated by the thermal 

degradation of the wood. Piloted ignition at heat fluxes sufficient to cause a direct-flaming ignition normally occurs at 

surface temperatures of 300 to 365 °C (Dietenberger 2004). 

Babrauskas (2001) summarized the findings of several 

works, spanning 30 researchers and more than 100 years, as 

they relate to the ignition temperature of wood. His findings 

show that all wood-based materials have similar ignition 

temperatures. 

When “sufficiently sized” wood members undergo exposure 

to elevated temperatures from a direct flame, thermal penetration occurs a distance inward from the base of the char 

layer. The immediately adjacent region of thermal degradation of the wood is referred to as the pyrolysis zone (Fig. 2). 

The char layer itself acts to insulate the wood. The exposed 

surfaces of sufficiently sized members decrease gradually. 

The charring rate depends on the species, orientation, and 

size of the specimen (Firmanti and others 2004). In the standard fire-resistance tests used in the building codes to regulate building elements, charring of wood occurs at a predictable rate for heavy timbers. The char rate of panel products 

and dimensional lumber will increase as the temperature at 

the center or backside of the wood product rise above the 

initial ambient temperature.

As wood burns, marked zones of degradation become apparent (if a specimen is thick enough). In response to a 

fully developed fire, the base of the char layer of the wood 

is commonly associated with a temperature of 300 °C. In 

Fahrenheit, a temperature of 550 °F (288 °C) has been used 

to identify the base of the char layer. Wood exposed to temperatures in excess of approximately 300 °C will form a 

residual char layer. Obviously, any charred portion of a fireexposed wood member has no residual load capacity. The 

wood beneath the char layer that is subject to some thermal 

degradation because of exposure to elevated temperature. 

This wood has residual load capacity but this residual capacity may be less than the load capacity prior to the fire. The 

normal wood (located beneath the zone of elevated temperatures) remains unaltered from the fire. 

Temperature gradients (associated with timbers) occur at the base of a char layer. The temperature gradient 

of elevated temperatures in uncharred wood starts at the 

base of the char layer (300 °C) and fades into undamaged 

wood (Fig. 3). In the standard fire resistance test, reported 

Figure 2—Degradation zones in a wood section. Figure 3—Temperature gradient beneath char layer in 

standard fire resistance test.
General Technical Report FPL–GTR–243

4

temperatures include 177 °C (350 °F) at 6 mm (0.2 in.) and 

104 °C (220 °F) at 13 mm (0.5 in.) beneath the char layer. 

Using the data of White and Nordheim (1992), Janssens and 

White (1994) determined the depth of the zone to be approximately 33 mm for large beams exposed in the standard 

fire resistance test. Beneath this layer is the normal or unaffected wood. There are equations and data for calculating 

char rates and temperature gradients that are valid when the 

member is thick enough to be considered a semi-infinite 

slab (White 2004). Additional information on thermal degradation, ignition, and charring of wood can be found in 

Browne (1958), White and Dietenberger (2001), and White 

and Nordheim (1992).

2.2 Elevated Temperature Impact on 

Mechanical Properties

Irreversible effects on the mechanical properties of wood 

can occur during the elevated temperatures associated with 

fire events. At a temperature of 66 °C (150 °F), the first 

loss of structural properties occurs, and significant degradation occurs within the temperature range of 200 °C to 

300 °C (392 °F to 572 °F), with the latter temperature being 

the base of the char layer (Ross and others 2005a). The irreversible effects of elevated temperatures on mechanical 

properties depend on moisture content, heating medium, 

temperature, exposure period, and the species and size of 

the piece involved (Green and others 1999; Kretschmann 

2010). The National Design Specification (NDS) for Wood 

Construction advises that prolonged heating to temperatures 

above 66 °C (150 °F) can cause a permanent loss of strength 

(American Wood Council 2012). Whereas days of heating at 

66 °C can have a permanent effect on mechanical properties, 

the immediate temperature effect on mechanical properties 

is reversible for heating periods that are limited to hours at 

temperatures below 100 °C (Kretschmann 2010). In contrast 

to the immediate impact on strength properties of individual 

wood members at elevated temperatures, the permanent 

loss in compressive strength occurs at higher temperatures 

than a comparative loss in tensile strength (Schaffer 1977, 

1982a,b). 

Effects of elevated temperatures are primarily on the 

strength properties. The effect on stiffness is considerably 

smaller. The NDS supplied multipliers that are applied to the 

tabulated stresses for wood that will experience sustained 

exposure to temperatures of 100 to 150 °F are more severe 

for bending and tensile stresses than they are for compressive or buckling stresses. In addition to temperature, adverse 

effects depend on duration and type of exposure. In contrast 

to the impact on strength properties of wood at elevated 

temperatures, the loss in tensile strength after cooling to 

room temperature is greater than the loss in compressive 

strength. While days of heating at 66 °C (150 °F) can have a 

permanent effect on mechanical properties, the temperature 

effect on mechanical properties is reversible for heating periods of hours at temperatures below 100 °C (212 °F). Thus, 

the zone beneath the char layer between 100 °C (212 °F) 

and 300 °C (550 °F) has the potential of irreversible loss 

in mechanical properties at varying levels due to thermal 

degradation. 

Without extinguishment, a fire has three phases in an interior structural fire: 1) the growth of the fire from ignition 

to flashover; 2) the fully developed post-flashover fire; and 

3) the decay period of declining temperatures as the fuel 

sources are consumed. Flashover is the full involvement of 

the combustible contents of the compartment and is associated with flames out the door in the standard room-corner 

test. Whereas temperatures are lower, presuming that the 

wood members self-extinguish, the cooling period of the 

decay period prolongs the duration of elevated temperature 

exposure. A prolonged cooling period is accompanied by 

diminishing surface temperatures, while the temperatures 

in the center portion of the cross section may still increase. 

Hence, elevated internal temperatures may result in further 

decline of residual properties. Additional information on 

elevated temperatures impact on mechanical properties can 

be found in Kretschmann (2010) and Schaffer (1984).

2.3 Fire Resistance Ratings of Exposed 

Wood Elements

Requirements for structural integrity during a fire are a major component of the building code provisions pertaining 

to fire safety. These requirements for buildings are in terms 

of fire resistance ratings of the structural components. The 

ratings depend on construction type and occupancy classifications. ASTM E 119 (ASTM 2010) describes the full-scale 

test methods used to determine fire resistance of structural 

members. Structural failure occurs when a member is no 

longer capable of supporting its designed load during the 

standard fire exposure. The fire exposure of the standard 

fire resistance test approximates the second phase, or postflashover portion, of the fire. Some of the specified temperatures for the standard time–temperature curve are 538 °C 

(1,000 °F) at 5 minutes, 843 °C (1,550 °F) at 30 minutes, 

and 927 °C (1,700 °F) at 1 hour.

The standardized fire test used to determine the fire resistance ratings of structural members has resulted in considerable data on the charring rate and temperature gradients of 

the remaining uncharred wood in a semi-infinite wood slab 

(Babrauskas 2001; Buchanan and Barber 1994) (Fig. 4). For 

such exposures, it is generally assumed that the temperature 

at the base of the char layer is 300 °C. As noted previously, 

the thickness of the wood layer with elevated temperature 

beneath the base of the char layer of elevated temperature 

is approximately 35 mm, and the temperature profile can be 

approximated by a parabolic curve. 

For exposed wood members, there are methodologies for 

calculating the residual load capacity during the standard 

fire resistance test and thereby determine the fire resistance 

rating of the wood member. The methodology with current 
Evaluating Fire-Damaged Components of Historic Covered Bridges

5

building code acceptance is described in Technical Report 

10 of the American Wood Council (2014) and includes information from the National Design Specification for Wood 

Construction (NDS) (American Wood Council 2012). The 

NDS specifies the design methodologies and property values to be used for the structural design of a wood structure. 

The equations used to estimate residual load capacity during fire exposure uses char rate equations to obtain reduced 

structural area of a member per time exposed to fire. One 

can account for additional thermal damage to uncharred 

wood in the load capacity calculations by including a zerostrength layer in the dimensions of the residual cross section. The NDS account for residual effects of wood in the 

zone of elevated temperature beneath the char layer as well 

as corner rounding by applying a multiplier factor of 20% 

to the char depth to obtain the reduced cross-sectional area 

of the wood (American Wood Council 2012). This reduced 

area is then used to solve for an estimated maximum capacity for a fire exposure time (American Wood Council 2014). 

The allowable design values in the NDS for the mechanical properties of a given grade and species are increased 

to their average ultimate strength values. Using the normal 

assumption of a char depth of 1.5 in. at one hour, the 20% 

calculation for the zero-strength layer is equal to 0.3 in. for 

a one-hour fire resistance rating. The methodology has been 

shown to be applicable to both dimension lumber members 

as well as heavy timber members.

2.4 Visual Grading of Lumber

From start to finish, wood construction incorporates numerous wood products into a number of primary and secondary 

structural applications. Wood is inherently variable and there 

are a number of factors that contribute to the performance 

of wood-based materials. Grading procedures account for 

the underlying factors of wood strength such as specific 

gravity, slope of grain, and the presence of knots. Grading 

procedures include visual grading criteria, nondestructive 

measurement such as flat-wise bending, stiffness, or density, 

or a combination thereof (American Wood Council 2012). In 

cases where graded lumber was used in the covered bridge, 

the grade stamp on the lumber may still be readable in areas 

protected from the weather.

For purposes related to post-fire assessments where the tabulated allowable design stresses for the grade are employed, 

it is recommended that the wood members be re-graded 

after the char is completely removed. The charring of the 

wood member is similar to ripping the wood member with a 

saw in terms of its impact on the mechanical properties and 

grade of the member. Visual stress grading rules are based 

on member size and characteristics (such as knots) of the 

outer zones. Thus, the removal of the outer zone as the result of the charring can greatly impact the stress grade of the 

member. Re-grading procedures take into account the impact of residual dimensions on the applicable grading rules 

for the reduced dimension as well as the altered relative locations of strength reducing characteristics in the cross section. To avoid confusion, the zero-strength layer could also 

be removed, or if it is to be retained as a protective surface, 

to account for it in the analysis. The re-grading of the members should be done by an appropriate a supervisory grading 

agency. The grading agency can make a qualified statement 

for each timber based on what grade characteristics are evident for each beam. A list of grading agencies is available 

on the web site of the American Lumber Standard Committee, Inc. (2015).

For the wood components that are not solid sawn lumber, 

the determination of the appropriate grade can be more difficult. Laminations of different grades are often used in construction of a glued-laminated timber. The inability to view 

the wide faces of the interior laminations will complicate 

the necessary re-grading of the residual beam. In addition, 

glued-laminated timbers use very high quality outer laminations and the loss of these laminations from a fire can severely affect the possibility of salvaging these timbers. 

2.5 Preservative-Treated Wood

Some wood in a covered bridge may be preservativetreated wood. While little data are available, the ignition 

and flammability behavior of water-borne preservatives are 

considered to be similar to untreated wood. Oil-borne preservatives, such as creosote, can be much more flammable 

than untreated wood. Weathering of creosote-treated wood 

appears to reduce the adverse impact of the creosote treatment. There is special concern for covered bridges carrying 

railroad tracks using creosote treated wood, as the potential 

for fire related to overheating and grinding of tracks from 

stuck brake mechanism is high.

One common preservative treatment for wood is CCA 

(chromated copper arsenate). A specific fire performance 

issue with CCA-treated wood is the potential for after-glow. 

Because of this after-glow behavior, fence posts have been 

known to be completely consumed hours or days after a 

Figure 4—Illustration of a charring wood member exposed 

to the standard fire exposure of 815 to 1,038 °C.
General Technical Report FPL–GTR–243

6

grass fire caused only minimal surface charring to the fence 

post before the grass fire self-extinguished. Because the 

continued after-glowing of the CCA wood is not easily visible, extra efforts need to be taken to insure that CCA-treated 

wood is truly extinguished. However, the fire risk posed 

by the CCA-treated wood may imply additional measures 

of fire protection than usual. The replacement of the CCAtreated wood itself will eliminate this concern.

For above-ground applications where leaching is not an 

issue, the boron-based preservatives are a common preservative. Boron is also a component of many interior fire-retardant treatments. Retentions levels for wood preservatives 

are much lower than those necessary for flame retardants.

3.0 Reduced Section Analysis of 

Fire-Damaged Wood Elements

The interior residual wood of timber stringers is protected 

from the intense heat of the fire by the char layer that develops. As a result, large structural wood members such as timbers do not necessarily need to be replaced after a fire (Buchanan 2001). Common practice with timbers (defined as 

those 5-in. by 5-in. and greater) neglects the reduced capacity associated with the elevated temperature zone. As such, 

once the char layer is removed, the remaining cross section 

is assumed to contribute fully to the load-carrying capabilities of the member (Firmanti and others 2004). Instances 

arise where negating the elevated temperature zone leads to 

un-conservative estimates of residual load-bearing capacity 

(Fig. 5). The wood beneath the char layer has residual load 

capacity; but this residual capacity may be less than the load 

capacity prior to the fire. A new cross section size is determined by carefully removing the char layer then measuring 

the char-free section dimensions. Research indicates that an 

additional reduction in the cross section may be warranted 

along with measurements of like members exposed to varying degrees of degradation (Kukay and others 2013).

Most available data on the structural performance of wood 

in a fire are based on research using the standard fire-resistance test. For the post-fire assessment, the exposure of the 

structural wood members to elevated temperatures during 

the decay period of fire development should be considered. 

Although the temperatures are lower during the decay period, the duration of the exposure can be prolonged compared 

with the duration of the fully developed post-flashover fire 

phase. The steep temperature gradient near the fire-exposed 

surface assumed in the normal assessment of residual load 

capacity during a fire is based on transient heating coupled 

with progressive charring of the wood cross section. During a prolonged cooling, the surface temperatures will 

decline while interior layer temperatures on the cool side 

may increase. Tests have indicated that this increase in the 

temperatures in the interior of the wood member caused by 

redistribution of the heat after fire exposure is particularly 

common for wood protected with gypsum board. Because 

the decay or post-extinguishment period is one of reduced 

temperatures, many observations of damage at the fire scene 

will be less helpful in establishing the intensity and duration 

of the exposure during this period. 

Recommendations for the additional reductions of cross 

sectional dimensions have varied. Often, it is assumed that 

the wood beneath the pyrolysis zone in a heavy timber 

structural member after a fire can be assumed to have full 

strength (Buchanan 2001). In Evaluation, Maintenance and 

Upgrading of Wood Structures, A Guide and Commentary

published by the American Society of Civil Engineers in 

1982, recommendations for evaluation of fire damage were 

for “removal” of a fixed amount of wood (Freas 1982). 

Recommendations included removal of the char layer plus 

approximately 1/4 in. or less of wood below the base of the 

char layer. For members controlled by compressive strength 

or stiffness, the recommendation was that no additional adjustment beyond removal of 1/4 in. was necessary to apply 

the basic allowable design stresses to the residual crosssectional area. For members controlled by bending strength 

or stiffness, the recommendation was either removal of an 

additional 0.625 in. or removal of an additional 1/4 in. in 

combination with a 10% reduction in the allowable design 

value used to calculate the load capacity of the residual 

cross-sectional area. 

In applying the methodology to the available strength data 

for permanent strength loss and the temperature profile reported for ASTM E 119 (ASTM 2010), White and Woeste 

(2013) concluded that 0.1 in. to 0.3 in. is a reasonable recommendation for the zero-strength layer of a member loaded in compression in a post-fire load capacity analysis when 

used with the NDS adjusted design values. For members 

loaded in tension or bending, the recommendations were a 

thickness of 0.3 in. to 0.5 in. These recommendations assume that the zero-strength layer is not physically removed 

Figure 5—Reduced section approach to fire-damaged wood.
Evaluating Fire-Damaged Components of Historic Covered Bridges

7

from the member and the temperature at the center of the 

timber did not increase based on the likely temperature profile during the fire. One can further adjust the depth of the 

zero-strength layer downward by a fraction (e.g., 50%) of 

any uncharred depth removed for appearance reasons. Selection of values between 0.3 in. and 0.5 in. should be based 

on the duration of the fire as reflected in the observed char 

depth and location of members relative to direct exposure 

to flames. The observed thickness of the residual char layer 

will be less than the observed reduction in the dimensions 

of the charred member due to shrinkage of the char layer. In 

the context of these recommendations, 0.3 in. for compressive members and 0.5 in. for tension and bending members 

are the more conservative values for the thickness of the 

zero-strength layer in the calculation of load capacity.

One approach is to use a series of assessments in applying the reduced section analysis to the fire-damaged wood 

(White and Woeste 2013). As a simpler assessment indicates 

the potential for retention of the damaged wood elements, 

subsequent analysis refines the input data and assumptions 

used to make the assessment. In the initial simple analysis, 

calculations are done using the following:

• Reference properties listed in the NDS for a common 

grade and species used for bridge construction. 

• Initial estimates of residual cross sections of the wood 

elements. 

• The charred layer is assumed to have no residual 

strength and stiffness but residual wood is assumed 

to be undamaged.

For those members identified as having the potential to be 

adequate for continued use, a more refined set of calculations can be made using the following:

• Reference properties listed in the NDS for the identified 

species of the structural element in the bridge being 

evaluated and its likely grade.

• More detailed measurements of the residual crosssectional areas and assumptions on the depth of an 

equivalent zero-strength layer can be refined.

Once the elements have reduced to their final dimensions, 

the final affirmative assessment of fire-damaged wood 

components needs to include visual grading of the residual 

structural elements. This re-grading needed for this final assessment would need to be done with the charred members 

cleaned of the char and reduced to their final dimensions. 

The maximum potential grade for each beam should be 

established and documented by an appropriate supervisory 

grading agency. Also, the responsible party should verify 

the final residual size of each timber for use when checking 

all fire-exposed timbers in the structure. The load history of 

the timbers may be important. The possibility of overloads 

in-service, or cumulative damage, should be investigated as 

well as damage from decay.

When dealing with calculations for residual load capacity of structural glued-laminated members, the structural 

grade variations of individual components that make up a 

composite wood member need to be considered. Gluedlaminated members usually have higher grades for balanced 

and layup beams that are appropriate for the outer tension 

and compression laminations than the interior laminations 

near the neutral axis; therefore, damage to the outer higher 

grade laminations can have a more significant impact on the 

residual load capacity of glued-laminated beams after a fire 

event (Ross and Pellerin 1994). The determination of the 

appropriate grade is also more difficult. The inability to visually inspect the wide faces of the interior laminations will 

complicate the necessary re-grading of the residual beam.

4.0 Post-Fire Inspection and 

Site Visit

A site visit is needed to assess the damage of a fire to a 

historic covered timber bridge. As part of this site visit, 

visual post-fire inspection process of individual members 

should facilitate an appropriate course of action. In so doing, technical personnel need to assess the residual structural 

integrity of individual members that exhibit varying degrees 

of degradation. The objective of the initial assessment is to 

identify those members that need to be replaced due to the 

fire damage and those members that warrant further investigation of their residual load capacity.

Observable variants can be documented using various pieces 

of equipment such as a moisture content meter (James 

1988), drill resistance press, increment core extractor and 

other equipment available for NDE assessment of wood 

members. Other issues related to decay and other pre-fire 

damage can also be noted at this time and analyzed. Prefire damage consists of degradation to a load-bearing solid 

wood member due in part to decay, insect damage, splitting 

or cracking, and overloading. Glued-laminated load-bearing 

beam damage can consist of the same degradations as 

solid wood, but also includes delamination and finger-joint 

defects (Garab and others 2010). All of these topics should 

be addressed during the post-fire evaluation.

4.1 Fire Itself

An understanding of the fire itself might aid in determining 

the degree of thermal damage beyond the base of the char 

layer. The standard for fire investigation is NFPA 921 Guide 

for Fire and Explosions Investigations (NFPA 2013), which 

can provide some guidance on the source of fire, the size of 

the fire, and its duration. Information gathered in the investigation of the fire itself may help establish likely maximum 

temperatures in various locations and the durations of the 

fire exposures. Although visual char depths may provide 

some insight, predicting a time of exposure to fire is difficult to determine from the depth of char and overall damage. 

Complicating factors including grain orientation, heat flux 
General Technical Report FPL–GTR–243

8

exposure, and moisture content can vary significantly, even 

within a given species. Accordingly, a uniform char rate 

cannot always be supported (Schroeder 1999). These caveats on wood charring should be taken into account when 

using the standard for fire investigations. To get further 

answers, professional fire investigators can be consulted, 

as they would have various investigative tools not available 

otherwise, such as chemical analysis of critical samples, fire 

modeling, and so on.

4.2 Decay Damage

While the inspection was initiated in response to a fire, an 

assessment of the damage should also verify that the timbers 

are otherwise “sound” and have not been damaged by factors other than fire exposure. Individual, biodegraded bridge 

components can be quantified with the use of NDE techniques to determine the impact of environmental conditions 

on the wood. Using NDE techniques, a complete analysis of 

the individual members can be determined without compromising structural integrity (Pellerin and others 1996). Many 

of the NDE discussed in this publication for potential use 

with fire-damaged wood are NDE techniques more widely 

used to detect decay or natural defects in structural member. 

The assessment of wood structures for decay is extensively 

discussed in the Wood and Timber Condition Assessment 

Manual (White and Ross 2014) and other publications.

The topics of biodeterioration of wood structures including 

bridges, options for in-place treatment to prevent or arrest 

the degradation, and the use of preservative treated wood in 

historic structures are discussed in other publications (Lebow and others 2012a, Lebow and Anthony 2012).

4.3 Moisture Content

A moisture meter is valuable for measuring the moisture 

content (MC) of wood members. Since the MC has a great 

effect on the strength properties of wood, the MC of individual members must be measured and recorded during the investigation. Guidelines for taking MC readings can be found 

in the Forest Products Laboratory (FPL) technical report, 

Electric Moisture Meters for Wood by William L. James 

(1988) and the Standard Methods for Use and Calibration 

of Hand-Held Moisture Meters, ASTM D 4444 (ASTM 

1992). Electrical resistance type meters are recommended 

along with 3-in.-long, insulated pin probes.

4.4 Dimensions of Residual Cross Sections

The initial site visit can be used to identify critical members 

that might warrant further investigation of the degree of 

damage to the mechanical properties of residual wood. For 

those members identified as having potential to be adequate 

for continued use, a more refined assessment involves more 

detailed measurements of the residual cross-sectional areas. 

The thickness of the char layer is likely different than the 

thickness of the wood charred as the result of shrinkage 

and surface recession. Thus, it is not advisable to derive 

the thickness of residual sections by subtracting visual char 

depths from original dimensions.

Beyond simple visual measurements of the dimensions of 

the residual cross sections, increment cores and resistance 

microdrilling provides means to document the dimensions 

of the residual core cross section, particularly where accurate visual measurements may be difficult.

4.5 Increment Cores

One option to examine the interior cross section of wood 

elements is to use an increment borer to extract a wood core 

(Ross and Pellerin 1994). Cross sections of the cores can be 

used to make density-related determinations. Commercial 

increment borers are often marketed for use on living trees. 

In addition to providing a visual measure of the thickness of 

the residual cross section, the core can be used to examine 

density gradient, depth of penetration of any chemical treatment, and presence of any voids caused by decay or insect 

damage. As will be discussed later, sections along the core 

can also be subjected to chemical assay or near infrared 

(NIR) analysis for evidence of thermal degradation. 

4.6 Drill Resistance Test

Whereas a simple hand or power drill can be used to attempt 

to determine the thickness of sound wood, the use of commercially available resistance microdrilling test equipment 

(Fig. 6) holds much promise. It is useful when qualifying the 

effects from decay, biodegradation, and fire damage. Past 

and current applications indicate that the drill-resistance test 

is useful when detecting the effects from decay and biodegradation (Ross and Pellerin 1994). By accurately denoting 

areas like this in individual members, inspectors can make 

a more informed decision on whether an individual member 

can remain in service (with or without repair), or if it should 

be replaced. Advantages to post-fire evaluations are that this 

technique may provide a direct qualifier of the residual capacity of individual members.

Modern equipment usually consists of an electric power 

drill, a specialized drill bit, and a software hardware interface to facilitate data collection and data analysis. Drill 

resistance tests are performed in much the same fashion 

Figure 6—Field test using commercial drill resistance test 

equipment.
Evaluating Fire-Damaged Components of Historic Covered Bridges

9

as drilling a pilot hole or a series of pilot holes into wood 

members. The results from a drill resistance test, however, 

are plotted as a function of insertion-depth zones that exhibit less resistance (Fig. 7) and can be noted for areas of 

uncharred wood.

Tests have been traditionally classified as a quasi-nondestructive (semi-destructive) test as minor localized damage 

is imposed on the member of interest during testing. Drillresistance tests are a function of drill-bit insertion depth. 

Generally speaking, a decrease in drill resistance or rather, 

an increase in the drill-penetration rate can be correlated to 

areas of decay, and aid in determination of the thickness of 

the residual shell (Emerson and others 1998).

Considering that current equipment can pick up on the 

changes in density from earlywood to latewood growth 

rings, a parallel may be drawn to that of the residual nondegraded wood. In the case of decay, the residual shell 

can be thought of as that portion of the cross-section that 

surrounds the deteriorated portion of wood; similarly, the 

residual shell is that portion of the cross-section that otherwise maintains load-bearing capacity. Conversely, for firedegraded wood, the residual capacity is the residual core of 

the member. This type of test has also been performed when 

assessing fire damaged concrete as documented in the works 

of Felicetti (2006).

Until recently little research had been applied to microdrilling to the evaluation of fire-damaged wood. In a quick assessment of the application of technology to fire-damaged 

wood (White and others 2013b) resistance microdrilling 

measurements were made on charred beam near the midlength of the beams in a vertical orientation (Fig. 7). The 

fire-charred outer portion of the glulam beam has considerably less density than sound wood and was easily detected 

with this NDE tool. The resulting relative density profiles 

were used to estimate (1) the remaining depth of uncharred 

wood and (2) the thickness of the thermal damaged wood 

layer. It was much more difficult to estimate the thickness of 

the thermal damaged wood layer.

5.0 Options for Evaluating 

Heat-Damaged Wood

The most obvious determination of damage from thermal 

degradation is visual. There can be various degrees of 

browning of the wood. Pyrolyzed wood retains the original 

cellular structure as observed via light microscopy or scanning electron microscopy (Zickler and others 2006). Zicherman and Williamson (1981) examined the microstructure 

of fire-damaged wood and found the demarcation between 

damaged and undamaged wood to be extremely narrow 

(several cell layers in thickness). Schroeder (1999) concurs 

with this observation of a narrow pyrolysis zone and attributed observations of a wider band to chemically borne moisture migration. For woods with resin contents, extrusion of 

the resin is evidence of exposure to elevated temperatures.

Charred wood is obvious. In some cases, there may be minimal or no char on the surface but smoke damage or surface 

resin is present to suggest some degree of fire exposure. In 

the reduced section approach, assumptions are made regarding the depth of the thermal degradation of the fire-exposed 

wood. Such assumptions are the basis for the thickness of 

the zero-strength layer that is added to charred wood to address the uncharred wood that has been thermally damaged 

and has reduced load capacity. The reality is that the fire 

exposure of a natural or real fire will be different than the 

fire exposure of the standard fire resistance test and there 

are considerable variations in such fire exposures. Assessment of light-frame elements and panel products that are 

not semi-infinite thick may also benefit from nondestructive 

evaluation (NDE) techniques to assess the degree of thermal 

damage. As previously discussed, an investigation of the fire 

severity and progression can provide some guidance. In this 

section, we discuss the potential NDE options for assessing 

thermal damage to wood.

While extensive work has been done on NDE techniques for 

assessing the residual capacity of wood affected by biological degradation, only limited work has been done on NDE 

Figure 7—Residual dimensioning using data from commercial drill resistance test equipment.
General Technical Report FPL–GTR–243

10

for wood exposed to elevated temperatures (White and others 2013a).

One exception was research done in the late 1980s and early 

1990s on NDE for the elevated temperature damage to fireretardant-treated (FRT) plywood roof sheathings (Ross and 

others 1991a; Winandy and others 1991a,b.) Certain FRTs 

accelerated the strength reduction degradation at elevated 

temperatures. NDE technologies have also been used in research projects to investigate damage from production processes, such as heat treatment and to investigate the chemistry of thermal degradation.

In addressing problems of thermally degraded FRT roof plywood, the National Association of Home Builders (NAHB 

1990) identifies five methods to determine the level of degradation in FRT plywood, including the following:

1. Concentrated proof loading

2. Removal of small samples for laboratory mechanical 

testing

3. Chemical analysis for chemical compositions of wood

4. Screw-withdrawal testing

5. Spectral analysis for end products of degradation

These methods are also potential methods for evaluating 

wood components for damage within the residual cross sections after a fire. As fire-damaged wood has a thermal gradient, removal of small samples for laboratory mechanical 

testing is probably not a practical option.

There are not standard or even commonly accepted methods 

for NDE of thermally degraded wood. A difficulty in evaluating wood for damage from elevated temperature exposure 

is the inherent variability in the properties of wood both in 

terms of the original strength properties and the properties 

being examined as an indicator of strength loss. Rather than 

predicting specific values of MOE and MOR, results for 

these techniques are likely to be better for comparing damaged members to similar members that are outside the area 

of elevated temperature exposure. From the comparative 

data, it may be possible to obtain evidence of either damage 

or the lack thereof.

As discussed, one detrimental outcome of the thermal degradation is the reduction in the density for wood exposed 

in temperature range of 200 to 350 °C. Taking into account 

both mass and volumetric changes, the change in density 

equates to 60% of original density at temperature of 340 °C 

(Zickler and others 2006). From 600 to 900 °C, the char 

density increases to about 80% of original density. Later at 

1,800 °C, the density drops back to 70% of original density 

(Zickler and others 2006). Thus, one option for NDE evaluation is to measure density itself. As discussed previously, 

two methods pertaining to density are the increment core 

and drill resistance techniques. One problem with using 

weight loss or reduced density as an indicator is that significant loss in strength can occur before there is significant 

loss in weight (Ross and Pellerin 1994).

For this discussion, the potential NDE methods have been 

divided into three groups. Methods in the first group are 

those that measure at or near the surface and include those 

of hardness, microindentions, and penetration as well screw 

withdrawal and ultrasonic. These methods reflect loss in 

density near the surface and resulting effects on properties. 

The second group includes methods that measure chemical 

changes in the wood associated with loss of residual load 

capacity. Besides chemical composition analysis, chemical methods also include radiography. Methods in the third 

group are (1) near infrared, (2) stress, (3) vibration, and 

(4) proof loading. These methods measure the overall interior properties of the wood element. These methods either 

directly or indirectly provide measurements of the modulus 

of elasticity of the wood.

Where individual members are considered, the techniques 

mentioned in this report are meant to supplement, not necessarily replace, existing approaches as they pertain to individual members that otherwise expected to remain in service.

5.1 Tests for Surface Material

Surface measurements can be taken to examine for damage. 

As these methods are actually measuring a small portion of 

wood, they are particularly sensitive to the inherent variability of wood. The results will be sensitive to the grain 

orientation in the area of measurements as well as moisture 

contents. Thus, multiple measurements are required to improve the validity of the evaluation. Moisture content can be 

quantified in the field with a moisture meter.

The simplest test for surface degradation is the “pick test.” 

It is a test used for detecting decay (Anderson and others 

2003). In the “pick test,” an ice pick or similar tool is used 

to poke the wood and ply out a splinter (parallel to grain) 

from it. With sound wood, the break will be some distance 

from the tool and from one end of the splinter. With degraded wood, the break in the wood will be at the location of the 

tool, and the brash, brittle break is perpendicular to the grain 

of the splinter. As with many of these tests, these subjective 

observations are best made by comparing results for the area 

of interests with results for similar wood elements in areas 

that are known to be undamaged. Although a simple technique, the subjective nature of the test makes it subject to 

misinterpretations. Two conditions that might lead to misinterpretation are water-softened wood and soft-textured wood 

(e.g., western red cedar) (White and Ross 2014).

The methods for the first group are those for hardness, 

indentation, and penetration. The methods for screw withdrawal and ultrasonic correspond to the second and third 

groups, respectively.
Evaluating Fire-Damaged Components of Historic Covered Bridges

11

5.1.1 Hardness, Microindentations, and 

Penetration Tests

These techniques based on hardness or penetration may 

provide an indicator of thermal degradation near the surface 

of the wood member. Zickler and others (2006) conducted 

nanoindentation tests on spruce wood that had been pyrolyzed to temperatures up to 2,400 °C and cooled to room 

temperatures for the indentation tests. Hardness increases 

from 0.4 GPa at the heat-treated temperature of 220 °C to 

approximately 4.5 GPa as treated at 700 °C and a reduction 

in values occurs starting with treatment at 2000 °C. The 

indentation ductility index remained constant at about 0.8 

until 300 °C treatment where it dropped to ~0.1 at around 

500 °C treatment. Zickler and others (2006) concluded that 

below 400 °C treatment, decomposition leads to a minimum 

of the elastic modulus. Above 400 °C treatment is a gradual 

transition from a visco-plastic biomaterial into a brittle, 

glass-like carbonaceous residue with maximum values at 

around 800 °C treatment. Above 2,000 °C treatment, plastic 

deformability improves.

Hardness test is the force required to indent a material a 

specified amount. There are various methods in wood hardness testing including the Janka method and the one specified by Branco and others (2009). The significant difference 

between the two methods is the size of ball indented into 

the surface of the wood. The Janka method uses a ball with 

a radius of 11.28 mm, and the aforementioned authors use a 

10-mm ball. The ball is indented into the wood to half of its 

diameter. The standard Janka method is a laboratory test of 

a specimen with specified dimensions. Portable commercial 

hardness devices exist but are largely intended for use on 

metal.

The microindentation technique is much like that of a traditional hardness test. A 1.5-mm (lengthwise), 0.5-mm diameter pin is pressed into the wood member with a very precise 

loading. Branco and others (2009) established a correlation 

between the force to indent the wood and modulus of elasticity in bending. This correlation was used in their experiment and would require more testing to be applicable in all 

cases (Branco and others 2009). Standard tests like this and 

the results can be found in Barletta (2005).

The amount of the pin that penetrates the wood has been 

used to generalize the mechanical properties of individual 

members. The penetration test, a variation of the microindentation test, uses a larger pin. This test measures the 

amount of strikes of a rebound hammer that is required to 

drive the large pin into the member a certain distance. The 

Pilodyn (KRS, Middelfart, Denmark) test uses the depth of 

penetration of a spring-loaded pin as a measure of degree of 

degradation (Ross and Pellerin 1994). Commercial Pilodyn 

devices are available.

Because very small localized compression is required to 

complete this test, the structural integrity of the member 

is not compromised during or after this test is completed. 

These tests would be run at the site of the member and could 

be completed relatively quickly. Given that the indentation 

is very small, many tests have to be taken to ensure that a 

proper sample of the properties of the wood are recorded. 

As a large portion of fire damage evaluation is measuring 

the residual strength for the member under the pyrolysis 

zone, this technique may not have the distance required to 

reach the zone of unaffected wood. A disadvantage is that 

the results may not indicate the initial strength losses because of the chemical changes in the early stages of thermal 

degradation.

5.1.2 Screw-Withdrawal Test

The screw-withdrawal test relates the maximum extraction 

load to residual flexural properties. Pilot holes are drilled 

into the undamaged end sections of each small specimens 

and a machine screw is then inserted so that it surfaces but 

does not extend beyond the specimen’s base. The machine 

screw is then extracted with a digital screw extractor. The 

screw withdrawal loads are recorded. The screw-withdrawal 

test is relatively fast to perform, and no major repairs need 

to be made afterwards.

The screw-withdrawal test was extensively investigated as a 

method for evaluating FRT plywood that had been thermally 

degraded in roof applications (NAHB 1990; Winandy and 

others 1998; Ross and others 1991a; White and Ross 2014). 

During the 1990s, this issue was a serious problem in the 

United States, so the portable commercial screw-withdrawal 

devices were available. Screw-withdrawal FRT plywood 

tests use a strain gauge attached to a metal collar placed 

over the head of a standard wood screw inserted into the 

underside of a piece of roof sheathing in the field. The screw 

is then withdrawn from the plywood using a tensile force, 

and the maximum load is recorded. The strength of the FRT 

plywood can be compared to the strength of untreated plywood by correlations between screw withdrawal force and 

the breaking force (Ross and others 1991a; Winandy and 

others 1998). In an initial FPL study of thermally degraded 

FRT plywood, regression of screw withdrawal resistance 

and bending strength had a correlation coefficient of 0.88 

(Ross and others 1991a).

Winandy and others (1998) inserted a No. 10 wood screw 

into various pieces of FRT plywood and untreated plywood 

that had been exposed to temperature greater than 130 °F for 

a specific time interval. They then measured the relationship 

between screw-withdrawal resistances versus the remaining 

plywood-bending strength and discovered that FRT plywood 

screw-withdrawal force is dependent on the thickness of 

plywood and various treatments methods. Based upon the 

method of fire-retardant treatment, the modulus of rupture 

was shifted either higher or lower. Models were developed 

for specific sub-groups, but data from different treatment/

thickness could not be grouped into a single universal model 

to predict the bending strength.
General Technical Report FPL–GTR–243

12

Kukay and others (2008) and Kukay and Todd (2009) developed equations including variables for moisture content, 

specific gravity, moment of inertia, maximum screw withdrawal load, cross-sectional orientation with respect to the 

pith, treatment group (charred or uncharred), and various 

interactions between these variables. They concluded that 

the results obtained through screw-withdrawal tests are best 

represented when the results are limited to comparisons to 

similar members that have obvious degrees of residual load 

capacity. The following reduced model was found to be adequate based on a model comparison F-test procedure:

Y = β0 + β 1MC + β 2SG + β 3I + β 4SW + β 5O

+ β 6T + β 7O • T + ε

The variables include moisture content (MC), specific gravity (SG), moment of inertia (I ), maximum screw withdrawal 

load (SW), cross-sectional orientation with respect to the 

pith (O), treatment group (T ) (charred or uncharred), and 

various interactions between these variables.

Models like this are believed to be applicable under similar field conditions and are methodology-, material-, and 

grade-specific. Rather than predicting specific values of Ef

and MOR, the results obtained through screw-withdrawal 

tests are best represented when the results are compared 

to similar members that have obvious degrees of residual 

load capacity. Variability of the results stem from changes 

in the predrilled pilot-hole size, the screw insertion depth, 

the screwtip to screw shank diameter, and the rate of extraction. For these reasons, care was needed when interpreting 

and extrapolating the results from screw-withdrawal tests. 

General correlations are likely to lack adequate precision to 

establish actual property values. The results of individual 

research studies that incorporate screw-withdrawal tests are 

generally not extrapolated. Additional work is needed to expand the models to account for the effects of a wider range 

of species and grades of materials as general correlations are 

likely to lack adequate precision to establish actual property 

values.

5.1.3 Ultrasonic Tests

Ultrasound technologies have been available since the 

1960s. Historically used for medical applications, they have 

been converted into a tool used to detect density and variation in density, decay, and/or defects near the surface in a 

variety of materials, including wood. Ultrasonic inspection 

can be used as a direct test to determine whether there is an 

immediate density change near the surface of the member in 

question. Ultrasonic inspection techniques consist of highfrequency stress waves that disperse quickly over a short 

distance in wood. Ultrasonic inspection has predominantly 

been used in manufacturing to estimate product quality, but 

can also be used to detect common strength-reducing defects such as knots, slope of grain, and decay. MOE is computed from the sound wave measurements. Measurements 

for MOE are only valid for longitudinal measurements but 

the velocity and frequency domain signal amplitude are useful for defect detection in the transverse direction (Klinkhachorn and others 1999).

Reinprecht and Panek (2012) have shown that ultrasound 

technology can be used to detect differential density in 

wooden members. In their study, sawdust was used to model 

rot in wood. The ultrasound waves that propagated through 

the media were significantly slower in comparison to a solid 

member. Density, or rather, specific gravity, is considered to 

be an underlying factor of wood’s strength; the greater the 

specific gravity, the greater the strength. The opposite could 

be said for members that show a decrease in specific gravity with regard to residual strength. Presuming that all other 

factors are the same, loss in specific gravity can be associated with wood members that have been exposed to elevated 

temperatures for an extended period of time. Klinkhachorn 

and others (1999) developed a portable ultrasonic device 

and conducted a few preliminary tests on charred yellow 

poplar specimens. There were little changes in the time delay or the velocity because of the charring of the two or four 

surfaces of the specimens but the area under Power Spectral 

Density (PSD) plots changed significantly as the result of 

the charring.

The position of the ultrasound probes can influence the 

results that are obtained from the ultrasonic inspection technique. The ability to obtain depthwise readings with in-situ 

technique like this becomes limited. Because of this limitation, some limit the technique to applications related to decay and other defects only (Emerson and others 1998).

5.2 Tests for Chemical Composition

Chemical analysis represents a different approach to NDE 

analysis of thermally degraded wood in that the effects being measured are the changes in the chemical composition 

of the wood. The changes in chemical composition are early 

indicators of thermal degradation and loss of residual load 

capacity. Small samples can be collected from the damaged 

wood and sent to a laboratory for chemical composition 

analysis. Alternatively, near infrared (NIR) technology can 

be used. In the investigation of options for assessing thermally degraded FRT plywood, the expense and specialization of testing were cited as disadvantages of the chemical 

analysis approach (NAHB 1990).

5.2.1 Chemical Analysis Tests

For initial exposures to heat, extractives and hemicelluloses 

are likely to be the first to be affected. Wood extractives 

such as fatty acids, fats, and waxes migrate to the surface 

of heat treated wood (Nuopponen and others 2004). Nuopponen and others (2004) used FTIR spectroscopy to examined Scots pine that had been heat treated under steam at 

temperatures of 100 to 240 °C. At temperatures of 100 to 

180 °C, resin acids in the radial resin canals moved to the 
Evaluating Fire-Damaged Components of Historic Covered Bridges

13

surface of the heat-treated wood and disappeared from the 

wood surface at higher temperatures (Nuopponen and others 

2004). Although not providing a measure of strength loss, 

consideration of extractive content can provide indication of 

the level of temperature exposure.

In the thermal degradation process, the hemicelluloses are 

among the first of the three main wood components to be affected (LeVan and Winandy 1990). Hemicellulose measurements have potential to provide direct measure of strength 

loss since the xylose, galactose, and arabinose content were 

sensitive to the loss of mechanical properties (Winandy 

2001). Arabinose was the most sensitive indicator of early 

strength loss. In a study on cyclic long-term temperature 

exposure at 82 °C, 30% RH, only arabinose showed a consistent reduction with increased durations of heating (Green 

and Evans 2008). Potential additional parameters that include holocellulose and alpha cellulose content and degree 

of polymerization has also been considered (Ross and others 

1991a).

Reductions in pH (increased acidity) of wood are a potential 

indicator of thermal degradation and strength loss. Lebow 

and Winandy (1999) investigated pH of wood as a technique 

for evaluating thermally degraded FRT plywood. To determine pH of plywood, Lebow and Winandy (1999) drilled a 

1/2-in.-diameter hole to a depth of 3/8 in. and collected the 

wood shavings to test the wood’s pH. After collecting the 

sample and mixing the sample cuttings with deionized water, they were able to measure the pH of the samples within 

20 min. Although some correlation (R2

 = 0.74) was observed 

between pH and strength loss in the plywood chemically 

treated with fire-retardant chemicals, this method was less 

sensitive to degradation in untreated plywood. Chemical 

analysis tests can be comparatively more expensive since 

most of the work has to be conducted in labs.

5.2.2 Near-Infrared Tests

The near-infrared (NIR) technology works by analyzing 

the interactions between materials and electromagnetic radiation. The wavelength and line speed are both variables 

predicted in models developed for detection of decay and 

other wood processing issues. Spectral analysis tests use 

infrared radiation to identify end products of chemical processes. A trained organic chemist is required to interpret the 

data gained from a spectrophotometer. More information 

on spectral analysis is provided in the American Plywood 

Association report, Fire-Retardant-Treated Plywood Roof 

Sheathing: In-Situ Testing (APA 1989).

The NIR technology can be used to accurately measure the 

chemical composition, mechanical properties, and a select 

few anatomical properties of wood (Brashaw and others 

2009). Esteves and Pereira (2008) have investigated NIR 

spectroscopy to evaluate properties of heat-treated wood. 

Additional information on this technique can also be found 

in the technical review by So and others (2004), Near Infrared Spectroscopy in the Forest Products Industry. Particular 

advancements have been made for the use of NIR technologies to detect the presence of in-situ member decay with the 

majority of studies being done specifically on brown-rot 

fungi (Nicholas and Crawford 2003). Other researchers have 

used this technique to predict surface moisture distribution as well as wet pockets. It is thought that wet pockets 

are associated with bacterial activity (Watanabe and others 

2010). The ability to detect local variants as well as general 

material distributive properties like moisture content make 

this technique appealing to applications beyond quality control. N is considered to be accurate and fast. More research 

is needed in the use of NIR technology to detect the early 

stages of degradation resulting in reduced load capacity. 

Considering its limitations, NIR is still a very attractive 

technology because it is lightweight and portable (Nicholas 

and Crawford 2003). These characteristics provide compelling reasons to further investigate its suitability for post-fire 

evaluation.

5.3 Tests for Whole Elements

For the interior characteristics of a wood element, the following measuring techniques are used. As useful tools for 

NDE assessments of wood decay, they are extended to thermally degraded wood elements. These include radiography, 

stress-wave, vibration, and proof loading. These techniques 

typically indicate changes in density or modulus of elasticity 

(MOE). Correlations are used to provide predictions for the 

modulus of rupture (MOR). Thus, the validities of the methods for predicting the residual load capacity based on MOE 

are limited by changes in the relationships between MOE 

and MOR caused by elevated temperature exposure. Stresswave techniques are widely used to assess damage from 

decay (White and Ross 2014). Proof loading techniques 

include those designed to assess the entire bridge construction. In situations where there is a progression of damaged 

members, destructive testing of selected members may be an 

option.

5.3.1 Radiography Tests

Radiography tests, also known as X-ray tests are carried 

out using an X-ray source and detector to establish the internal density distribution of wood members (Niemz 2012). 

The technique typically involves either two-dimensional or 

three-dimensional imagery and is dependent on the equipment and access points. This technique, like many others, 

can quantify the extent of decay as it is correlated with density losses. Since density losses from thermal degradation 

are also correlated to losses in flexural properties, radiography technique is applicable to post-fire investigations. In 

particular, radiography can detect a redistribution of density 

in cases where bulk density does not change because of liquid solidification or volatile condensation within the wood 

bulk. Indeed, radiography is commonly used for detecting 
General Technical Report FPL–GTR–243

14

internal voids in wood mostly during manufacturing (Ross 

and others 2005b).

It has also been used to determine variants such as knots as 

well as the internal decay of individual wood members in 

laboratory settings (Niemz 2012). A technique such as this 

could currently apply to in-situ depictions of the pyrolysis 

zone of individual wood members.

X-ray technology is able to pick up the presence of knots 

or defects based on density changes in the members. This 

would have the potential to allow inferences to be made 

about the residual strength of individual members beyond 

simply reducing the cross-section and visual inspections.

Radiography could be used in the field for in-situ testing if 

further advancements are made in making the equipment 

more mobile. High costs of equipment, inspector safety factors, portability, member access, and expert interpretation of 

radiograph results are chiefly responsible for keeping X-ray 

techniques from being more widely used as a field-evaluation technique (Ross and others 2005b). As a result, the 

technique may not have advanced far enough at this point.

5.3.2 Stress-Wave Tests

Stress-wave evaluation has been successful in detecting 

decay in structures (White and Ross 2014; Brashaw and others 2004). Stress-wave evaluation is based on sound-wave 

physics. Stress waves (also referred to as sound waves) are 

generated from an impact on the surface of a member that 

is under investigation. Impact sound waves are measured as 

a function of their time to travel from one side of the member to the other side. Measurements are made so any and 

all regions of decay are captured (White and Ross 2014). A 

guide specific to stress-wave evaluation is Stress-Wave Timing Nondestructive Evaluation Tools for Inspecting Historic 

Structures by Ross and others (2000). The methodology is 

also extensively discussed in the Wood and Timber Condition Assessment Manual (White and Ross 2014). Sound 

waves have a lower velocity through decayed wood than 

through non-decayed wood. Knowing the member’s cross 

section dimensions, an inspector can successfully locate 

areas of decay by making a series of measurements at incremental locations along the member (Emerson and others 

1998). Stress-wave evaluation tests have more recently been 

used as indicators of residual wood stiffness for individual 

members.

The handheld device used to perform the stress-wave analysis is portable, which makes it a viable option for essentially 

any location. The test is non-invasive and very easy to 

complete in a field setting. The test is efficient and can run 

successive times in a timely manner. The moisture content 

of the member must be known before the analysis can be 

completed, which can also be found using a portable moisture content meter. Because this test relies on stress waves 

quickly propagating through the member, it is essential that 

the exact distance between the first and second probes is 

known. Any deviation in the measurements between readings would make the analysis highly unreliable because of 

the high-speed travel of the stress wave.

Impact sound waves are measured as a function of their 

time to travel from one side of the member to the other 

side. Dynamic modulus of elasticity or apparent MOE (Pa) 

is considered equal to the wave speed (m/s) squared times 

gross density (kg/m3

) (Ross and Pellerin 1994). As with 

some other NDE methods, stress wave is a better predictor 

of MOE than it is for MOR. Elevated temperature exposure 

changes the correlation between the MOE and the modulus 

of rupture (MOR). Thus, a disadvantage for NDE based on 

determining elastic properties is that the property of most 

interest in terms of safety (i.e., strength properties) are more 

sensitive to elevated temperature exposure than the properties being measured.

Garcia and others (2012) used stress wave for nondestructive determination of the MOE of wood before and after 

heat treatment. The dynamic modulus of elasticity decreased 

by about 13% in the most severe treatment (230 °C for 4 h) 

but not for the milder treatments. A strong relationship was 

found between energy absorption as measured by stresswave techniques and residual strength of degraded FRT 

wood but the speed of sound was insensitive to the degradation (Ross and Pellerin 1994).

As part of our study on evaluating fire-damaged components of historic covered bridges, four 127-mm by 305-mm 

beams were subjected to a fire exposure in a fire test furnace that resulted in a narrow char layer but a deep layer of 

elevated temperatures that penetrated beyond the base of 

the char layer by nearly 25 mm. As a result of the fire exposure, average measured values indicate a 54% reduction 

in the maximum recorded load, a 46% reduction in strength 

(MOR) based on the residual cross-sectional area, and a corresponding 20% reduction in stiffness (MOE). These results 

are based on a standard three-point bending test with an increasing load applied at the lengthwise center of each beam. 

Sound wave measurements were taken near the top and 

bottom of each beam. The values and corresponding results 

were repeatable; resulting comparisons of the transmission 

rates indicated a reduction from the fire exposure. But the 

reductions in the times for fire exposure were not greater 

than the differences between the fastest and slowest times 

for the different specimens prior to fire exposure.

5.3.3 Vibration Tests

Vibration tests are conducted by placing an electric motor 

with a weighted wheel (attached) into motion at the midspan of a bridge or another readily identifiable location. This 

causes the bridge to resonate. Piezoelectric accelerometers 

are then placed and are used to measure the response to the 
Evaluating Fire-Damaged Components of Historic Covered Bridges

15

vibration (Wang and others 2005). In so doing, a number of 

modes of vibration can then be recorded. These recordings 

are used to assimilate computer-generated models that are 

calibrated to individual bridges.

Typically, the nondestructive forced vibration method has 

been used to determine the modulus of elasticity of structural members and structural systems. When a number of 

individual members are targeted, the stiffness of a member 

or rather a decrease in stiffness may be used for evaluation 

purposes. An empirical correlation between stiffness and 

strength can be used to determine the soundness of a particular wood member (Emerson and others 1998).

Vibration tests continue to be employed when establishing the in-situ dynamic behavior of pre-stressed, pre-cast 

concrete bridges, steel bridges, and composite bridges. Depending on the goals, it is possible for vibration tests to be 

conducted on individual members after being removed from 

service.

The advantage of vibration tests such as this are that with 

further research it could be used to quickly determine the 

stiffness of a timber bridge as well (Wang and others 2005).

Whereas forced vibration tests can be used to assess the 

dynamic behavior, more research is needed in the area of 

post-fire assessment using current methods (Wang and 

others 2005). Also, the weight of the structure needs to be 

accurately determined to predict bridge stiffness based on 

the beam theory model (Wang and others 2005). Estimating 

the weight of a bridge structure is difficult and may be construed as unreliable.

Dynamic testing and diagnostic load testing are two of 

the most common methods used for assessing the overall 

condition of a bridge. The response to the dynamic system 

load can be evaluated against an analytical model of the 

bridge or, if available, a previous response record. Vibration 

analysis techniques have been demonstrated on simple-span 

members but are more complex when full-scale structures 

are under investigation because of the structures’ many 

modes of vibration. Each mode of vibration must be investigated to determine the structural characteristics of the bridge 

(Emerson and others 1998).

5.3.4 Proof Loading Tests

Mechanical testing is a more direct approach to evaluating 

the effect of elevated temperatures on mechanical properties. In terms of NDE, this normally involves some sort of 

proof loading to determine the MOE. A bending proof load 

test was one of the options developed for the assessment of 

FRT plywood roof sheathing (NAHB 1990; Ross and Pellerin 1994). As part of the study on fire-damaged components 

of historic covered bridges, a proof loading method specific 

to field application to fire-damaged wood components was 

developed (Kukay and others 2013).

As part of this project on NDE of fire-damaged wood, new 

NDE proof-load techniques were developed (Appendix).

Dynamic testing and diagnostic load testing are two of the 

most common methods used for assessing the overall condition of a bridge. To directly determine a structure’s load-carrying capacity, one approach requires that large proof loads 

be applied to the bridge. These large proof loads usually involve rolling heavy vehicles onto the bridge while measuring strain or deflections of various members. The measured 

strain data is then analyzed to establish the strength of the 

members. Although smaller static loads have also been used 

for this purpose (Emerson and others 1998), determining the 

maximum load-carrying capacity of the structure is generally not as reliable as when larger proof loads are used. This 

requires an assumption (or sampling) for the material properties to be reliable.

5.3.5 Destructive Mechanical Tests

Mechanical testing options include destructive methods 

such as removal of selected members for testing or removal 

of small samples for laboratory mechanical testing. Destructive testing will provide a direct measurement of modulus of 

rupture while proof-loading techniques only provide direct 

measurements of modulus of elasticity. The removal of 

small samples for laboratory mechanical testing is well suited for both the field and the laboratory. The span-to-depth 

ratios, based on uncharred dimensions, dictate the amount 

of material that must be removed for mechanical testing. A 

minimum span-to-depth ratio of 14:1 is common for small 

clear samples of sawn lumber products and is specified in 

the standard ASTM D 134 (ASTM 1994). The standard 

ASTM D 198 (ASTM 2009b) specifies a span-to-depth ratio 

between 17 and 21 that is typically used for bending tests 

of lumber in structural sizes, while 18:1 frequently serves 

as the international standard for wood products specified in 

standard ASTM D 3043 (ASTM 2000). Depth-wise reductions from charring would otherwise increase the span-todepth ratios accordingly.

In accordance with ASTM D 143-94 (ASTM 1994), a 

single-point loading scheme can be applied at mid-span 

for all simply supported small members. Displacement 

measurements are taken during the test at the lengthwise 

midpoint, where the load is applied continuously to failure 

at an acceptable rate of motion. Flexural properties are then 

determined by analyzing the data that are collected from 

such tests.

Bending load strength tests are based on Test Method A that 

is specified in the standard ASTM D 3043 (ASTM 2000). 

Either 1- or 2-in. pieces are used for panel thicknesses less 

than 1/4 in. Two-inch pieces are used for greater thicknesses. The length of each test specimen depends upon the 

orientation of the laminations to the span. Perpendicular 

lamination test specimens shall not have a length of less 
General Technical Report FPL–GTR–243

16

than 24 times the depth plus 2 in. Parallel laminations test 

specimens shall not have a length of less than 48 times the 

depth plus 2 in. It is recommended to take at least three 

samples per panel.

Test samples shall be tested with a continuously prescribed 

loading rate—without shock—until failure to determine 

modulus of rupture. Measurements of deflection can be taken simultaneously to calculate modulus of elasticity, but one 

could be better served to use (Test Method B that is specified in the standard ASTM D 3043 (ASTM 2000)) instead 

for modulus of elasticity testing.

Bending load tests are fairly easy to perform with the correct equipment. Low costs are associated with the bending 

load tests, but the cost to repair the bridge must also be considered along with the shipping costs.

“The strength properties of various species are determined 

through standardized tests of small, clear (defect free) specimens at various moisture conditions. Large pieces, containing defects typical of standard grades of lumber, are also 

tested to develop strength data” and correlation between 

small clear specimen and standard structural sizes typically 

assessed from test method ASTM D 198 (ASTM 2009b). As 

addressed in the American Society for Testing and Materials, ASTM D 143-94 (ASTM 1994) Standard Test Methods 

for Small Clear Specimens of Timber, “The need to classify 

wood species by evaluating the physical and mechanical 

properties of small clear specimens has always existed”.

This technique presents one of five post-fire assessment options that were originally proposed by NAHB for directly 

determining the residual load-bearing capacity of thermally 

degraded plywood (Winandy and others 1998) across a 

broader range of materials. Similar approaches have been 

adopted in the works of LeVan and Winandy 1990 in their 

study on the effects of fire-retardant chemicals on bending 

properties of wood at elevated temps and Glos and Henrici 

(1991) when they assessed the bending strength and bending 

stiffness of structural timbers at temperatures up to 150 °C.

Advantages to post-fire evaluations that make use of this approach are such that much of parent material remains. This 

allows for the fire-damaged member to be sistered onto, 

and may otherwise eliminate the need for a temporary support mechanism, depending on the severity. This presumes 

that the primary connections are otherwise determined to be 

adequate. In instances where a fire is contained, the residual 

flexural values of like materials that exhibit varying degrees 

of degradation can be directly compared to one another with 

this technique.

Disadvantages are such that the removal of a smaller member from an existing member requires repair, if it is otherwise expected to remain in service.

6.0 Other Post-Fire Assessments

6.1 Fire-Damaged Connections

During a fire, heat can also be transferred into the wood 

through exposed metal connections (Fuller and others 

1992). Structural integrity of the covered bridge’s truss is 

heavily dependent on metal connectors. Therefore, all connections require a detailed post-fire inspection to determine 

their load-bearing capacity. Connection integrity is dependent on the quality of the metal and metal surface contact 

area. Chemical damage from corrosive effects of fire residues that may be present must be accounted for (Williamson 

1982a,b). Figure 8 illustrates a bearing style connection. 

Because metal can conduct heat into the interior of the wood 

component, assessment of fire damage around connections 

may be less feasible for visual inspection. Radiography 

tests, or also known as X-ray tests, may be advantageous for 

such assessment.

Inspections of the connections should also include examinations for corrosion of the metal fasteners. The surrounding 

wood should be examined for splits, decay, and insect damage. The tightness of the connections should be assessed.

6.2 Smoke Damage

Besides the stated effects of fire on wood (charring, reduced 

load-bearing capacity, and fire-damaged connections), 

smoke and other fire residuals will also be present; however, 

these do not affect the wood members’ load capacity (King 

2002). Nevertheless, sealing and eliminating fire odors is 

still important. Paints, sealers, and other finishes are available to seal the odors in the member material. Before the 

application of these sealers, the wood member should be 

thoroughly cleaned and exhibit no traces of the previous 

fire damage. This is true even for members that will be hidden from sight. A post-fire restored historic covered bridge 

should not only look as it did pre-fire, it must also exhibit 

no lingering fire damage odor. The National Institute for Disaster Restoration (NIDR) has guidelines for fire and smoke 

repair. These guidelines highlight methods for removal of 

fire residues, removing fire odors, neutralizing acid residues, and the use of sealing and encapsulation.

Figure 8—Bearing style connection, Red Covered Bridge.
Evaluating Fire-Damaged Components of Historic Covered Bridges

17

7.0 Post-Fire Bridge Repair

7.1 Historic Covered Bridge Rehabilitation 

Guidelines

Wacker and Duwadi (2008) state that when undergoing historic covered bridge rehabilitation, strengthening methods 

should conform to the Secretary of Interior’s guidelines 

for preservation of historic structures. In that regard, any 

rehabilitation necessary to increase or improve the strength 

of the structure should be done in a way that the historical 

external appearance is preserved (Figs. 9 and 10).

According to the Vermont Historic Bridge Program (2014), 

a historic covered bridge restoration needs to ensure the 

maximum load capacity associated with the anticipated use 

of the bridge is maintained in accordance with a prescribed 

set of preservation guidelines. This methodology will ensure 

that rehabilitated historic covered bridges will maintain their 

historic status and charm, but will also meet traffic and environmental demands. 

The Vermont Historic Bridge Program (2014) issued the 

following set of rehabilitation guidelines in order of priority, with the highest priority listed first and decreasing 

thereafter:

1. Retain all existing historic materials that have not deteriorated beyond the point of repair. Where existing 

rot or other damage is not severe enough to require replacement, the materials should be repaired rather than 

replaced. This treatment should be applied to each member individually, and deterioration of a large number of 

bridge elements should never justify the replacement of 

any single member capable of being replaced.

2. Replacement of existing materials in kind, meaning 

identical species, quality, and dimension to the maximum extent feasible, or restoration of original materials 

and design. Preferably, material origins should be from 

the same region of the specimen. If a different species 

or quality is considered and/or materials from the same 

region of the country are not available, substitutions may 

be considered with justification.

3. Application of historic methods of strengthening such 

as the application of sister lattices in Town lattice truss 

bridges.

4. Introduction of glued-laminated beams as a co-functional, reversible structural system. The beams must be designed to work in conjunction with the historic structural 

system to achieve required load capacity, and the historic 

structural system must be restored according the Rehabilitation Guidelines 1, 2, and 3 above.

Figure 10—Underside of bridge deck, Red Covered Bridge.

Figure 9—Captain Swift Bridge near Princeton, Illinois.
General Technical Report FPL–GTR–243

18

5. Replacements of limited pieces of existing load-bearing 

members with materials identical in species, quality, and 

origin to the maximum extent feasible. Dimensions may 

be larger but must not cause alterations to the dimensions of any other important bridge components. For 

example, increasing the depth of bottom cords of Town 

lattice trusses may increase capacity without requiring 

alteration to either overall bridge dimension or the design of the floor system.

6. Replacement of existing load-bearing members with 

glued-laminated members (beams or chords) of identical 

dimension.

7. Reinforcement of load-bearing members with nonobtrusive modern materials such as steel rods or plates, 

glass fiber, carbon plates, or other material. Placing these 

modern materials hidden from view, to maintain the historic integrity of the bridge, whenever possible. 

8. Protection of load-bearing members by the introduction 

of steel beams that provide a safety-net for the bridge. 

The redundant structure must allow the existing timber 

frame to continue functioning, and a minimum clearance between steel beams and floor beams should be 

designed. The purpose of this treatment is to protect the 

historic bridge in case of structural failure, not to increase carrying capacity.

9. Replacement of load-bearing members with, in order 

of priority: (a) timber of larger dimension but otherwise 

identical in terms of species and quality; or (b) timber of 

larger dimension and different species.

10. Replacement of existing load-bearing members with 

modern materials.

These guidelines may help post-fire investigators, bridge 

engineers, and bridge owners to develop a rehabilitation 

strategy that incorporates appropriate guidelines and requirements. With a consistent approach, case studies can be 

conducted to further assist an engineer with any problems or 

questions that occur during their rehabilitation efforts.

7.2 Historically Accurate Reinforcement and 

Rehabilitation

Conventional strengthening/reinforcing methods for timber 

structures may be accomplished through the use of member 

augmentation such as steel plates or bars, aluminum plates, 

or simply timber patches. On structures where it is impractical or impossible to replace a member, member augmentation may be a method that is capable of replenishing the 

load capacity of a damaged member. Sistering and splicing 

are the two most commonly used methods of member augmentation. When a member’s existing capacity is insufficient, sistering a member by adding reinforcing pieces either 

over a portion or even the entire member is the most popular 

choice. Sistering may also be used to repair members that 

are exposed to weathering or exhibit significant changes in 

moisture content as compared to splicing (Ritter 1992).

The sistered members can be attached to the damaged member in a number of ways. Two approaches include either adhesive bonding or steel plates with bolts. Adhesive bonding 

can have several advantages over steel plates: the stress can 

be more uniformly dispersed throughout the member, little 

weight is added to the structure, little to no damage occurs 

to the adherents, and it is generally less expensive. Of all 

adhesives for in-situ repair, epoxies have demonstrated minimal shrinkage during curing, remain dimensionally stable 

after hardening, and maintain a high resistance to water. 

After a surface has been cleaned, the epoxy augmentation 

process should begin immediately. A mechanically sound 

adhesion can be produced with relatively little force being 

applied to the scabbed member during curing (Custodio and 

others 2009).

7.3 Modern Reinforcement and Rehabilitation

Conventional repair methods may increase dead loads, installation costs, and transportation expenses because heavier 

reinforcement materials are used. Modern reinforcement 

methods using fiber reinforced polymers (FRP) have lower 

strength-to-weight and stiffness ratios. This allows FRPs to 

be used without a significant increase to the dead load of a 

structure.

FRP composites are usually bonded to the higher stress 

zone, more commonly known as the tension side of timber 

beams, which increases their load-carrying capabilities 

while decreasing deflection. FRP composites bonded on the 

tension face are either adhered to the surface as a sheet or 

are inserted into a cut narrow groove and secured into place 

with a bonding agent. Near-surface mounting of FRP bars 

placed in cut grooves is a new technology designed to increase the load-carrying capacity and energy absorption capacity when compared to unstrengthened counterparts (Kim 

and Harries 2010). Installing the bars in a cut groove on the 

tension side of the members means that repairs will typically be hidden from view. This is very important to historic 

covered bridges that need to maintain their historic integrity 

and charm.

Glass-fiber reinforced plastic (GFRP) bars are generally less 

expensive than other FRP bars, carbon fiber and aramid fiber, but still significantly improving a member’s load-carrying capacity and stiffness (Raftery and others 2009). When 

the Tourand Creek Bridge was strengthened with GFRP 

bars, it cost less than 15% of the estimated cost to replace 

the entire bridge ($800,000) (ISIS Canada 2000).

GFRP bars have a high strength-to-weight ratio and are noncorrosive. Both of these attributes provide engineers with 

a secure long-term option to reinforce in-situ members of 

historic timber bridges.(Gentile and others 1999) installed 

GFRP bars into grooves cut into the tension side of select 
Evaluating Fire-Damaged Components of Historic Covered Bridges

19

structural Douglas-fir stringers. The GFRP bars were bonded in place with an epoxy resin and the reinforced stringers were tested against similar non-reinforced stringers. 

Results showed a strength increase of 25% to 50% for the 

GFRP members. Ultimately, the study concluded that GFRP 

bars are a viable treatment method to increase the flexural 

strength of sawn timber beams.

The use of epoxy adhesives to bond the FRP bars into place 

has increased more recently based on their gap-filling qualities and low clamping pressures (Raftery and others 2009). 

To prepare the wood surface prior to bonding a FRP, it must 

be planed where bonding is going to occur; adequate bonding has to occur; degreasing of the surface has to be done; 

and the surface also has to be free of loose material (DOT 

2007).

The U.S. Department of Transportation conducted a largescale bending test with GFRP plates that were bonded to the 

tension face of several timber members. This research concluded that the strength and stiffness of timber members was 

vastly improved (DOT 2007).

Aramid fibers are another FRPs option that consists of a 

synthetic organic polymer fiber. Yahyaei-Moayyed and 

Taheri (2011) found that these fibers exhibit the highest 

tensile strength-to-weight ratio of all reinforcing fibers and 

are mainly used in industry, military, and aerospace applications. They also demonstrated that aramid fibers can successfully be used to increase the flexural strengthening of 

southern yellow pine and Douglas-fir species by 74% and 

34%, respectively.

Another method of strengthening wood beams is to bond 

steel cords to the beam’s tension face. Borri and Corradi 

(2011) demonstrated the effectiveness of bonding highstrength steel filaments to the tension face of wood beams 

by simply placing steel filaments on the wood surface and 

covering them with a thin layer of epoxy.

For all cases of FRP bar installed into timber, the bond 

shear-slip between the two materials has a minor effect 

on the ultimate loading capacity of the members. During 

analysis, a continuous bond between the two materials was 

observed in most of the tested cases (Valipour and Crews 

2011).

7.4 Member Replacement

After a post-fire analysis of the damage of the structure is 

complete and residual properties of the materials are determined, it may be necessary to decide whether to replace 

or to reinforce and rehabilitate a damaged member. More 

specifically, structural members that do not meet current 

load requirements and applicable codes should be repaired, 

reinforced, or replaced. Selected members should be analyzed based on which repair method is the most feasible, 

while maintaining historic covered bridge repair principles 

and guidelines. A list of these guidelines is presented in the 

next section. Abrasive blasting may be employed to remove 

char and member discoloration. Various media types can be 

used including ground corn cob, sand, and baking soda. The 

final wood surface should be treated for residual odors, and 

protective sealers should be applied (White and Ross 2014). 

If the member is damaged beyond repair, replacement is 

typically the final option. For historic covered bridges, 

replacement should be identical down to its cross section, 

species, and harvest location to retain the bridge’s historical 

integrity. When this option is not available, glued-laminated 

members may be used in accordance with the applicable 

guidelines (see APA 2009). Rodgers and others (1997) 

showed that replacing existing solid-sawn stringers that 

had accumulated more than 50 years of service with new 

glued-laminated stringers reduced the overall deflection of 

the structure under load and made its load-bearing response 

more uniform.

8.0 Design for Fire Prevention and 

Control

Minimizing fire damage to wood members is of great importance to the rehabilitation, repair, and restoration of 

historic covered bridges. Post-fire considerations should 

include an assessment of actions that can be taken to reduce 

the likelihood and severity of future damaging fires. With 

their construction and often isolated locations, covered 

bridges are a challenge for fire protection engineers. Riskbased strategies associated with fire containment and suppression through fire-retardant treatments are intended to 

prevent a bridge from collapsing in the event of a fire, but 

does not necessarily prevent a fire from occurring (Barbaresi and others 2012). Approaches like this should also be 

considered to be part of the design. Site-specific conditions 

and proximity to resources necessitate a combination of effective techniques into the design and/or retrofit. Design 

considerations related to fire prevention and extinguishment 

programs should incorporate serviceability, preservation, 

and aesthetics.

Although arson is a major cause of fire ignition to covered 

bridges, the possibility of naturally occurring and unintentional fires near a covered bridge is possible near the wildland–urban interface. Diligent removal of ignitable materials 

can curtail a fire’s preferential pathway. Thus, one obvious 

consideration should be the improved maintenance to reduce 

contents or surrounding vegetation that can add to fire intensity or ease of ignition.

Whereas it is beneficial to remove surrounding vegetation or 

other combustible items that provide a path for a wildfire to 

spread to the covered bridge, it needs to be recognized that a 

major avenue for the spread of wildfires to structures is the 

firebrands and embers lofting considerable distances from 

the wildfire itself.
General Technical Report FPL–GTR–243

20

In situations where elements of the bridges are being replaced, consideration of the relative fire performance of 

different building materials should be a factor in the selection of materials. The flame spread behavior of wood is 

dependent on the species. Chemical treatments are available 

to further reduce the rate of flames spread when the wood 

is ignited. Larger structural elements are less conducive for 

ignition and sustained flaming. Selected installation of panel 

products such as gypsum board can be used to reduce the 

areas of potential fire spread into smaller compartments.

Historic covered timber bridges are often in remote locations. Fire on these structures will likely be reported after 

the fire is fully developed or even after the decay period of a 

fire. Remote security and fire detection monitoring systems 

for these historic covered timber bridges can be installed to 

reduce response times, preferably to where only a small or 

no fire damage occurs. Design considerations that accommodate the retrofit of fire notification and extinguishment 

systems should include routine operation and maintenance 

inspections; both of which can be reduced through tamperresistant installations among other factors. The isolated locations of many historic covered bridges make some options 

for improved security topics of debatable benefits in terms 

of vandalism. Improved lightening generally will deter vandalism but in remote locations it may facilitate the unauthorized use of the covered bridge as a gathering place.

8.1 Flame Spread Index of Wood

Treatments are available to reduce the flammability of lumber and other wood products. In the building codes, regulations pertaining to surface flammability of interior finish are 

based on the surface burning characteristics test described 

in ASTM E 84 (ASTM 2009a). This test is widely known as 

the 25-ft tunnel test. In the test, observations of flame travel 

along the specimen are used to calculate a dimensionless 

flame spread index (FSI). The standard ASTM E 84 test has 

a duration of 10 minutes or less. Historically, the index was 

equivalent to zero for asbestos and 100 for red oak flooring. 

Measurements of the obscuration of a light beam are used to 

calculate a dimensionless smoke development index (SDI) 

which has a value of 100 when red oak flooring is tested. In 

the building codes, the provisions pertaining to the fire ratings of interior finish are in terms of the three classes of A, 

B, and C. Class A is the most restrictive and requires a FSI 

of 25 or less. For wood products, Class A classification typically requires a fire-retardant treatment. Class B requires a 

FSI of 75 or less (i.e., 26 to 75). For lumber, some domestic 

softwood species and some high-density imported hardwood 

species have reported FSI in the Class B range. Class C is 

a FSI of 76 to 200. Most wood panel products, lumber of 

some domestic softwood species, and the lumber of the 

domestic hardwood species have reported FSI in the Class 

C range. Lumber is typically tested in the E84 test with a 

nominal thickness of 1 in. (actual thickness of 19 mm). The 

FSI may increase with reductions in the thickness of the test 

specimen. Some wood panel products have been reported 

with FSI that exceed the upper limit of 200 for Class C. The 

requirement for the SDI is typically 450 for all three classifications. Available SDI data for wood products are for SDI 

less than this 450 requirement. The American Wood Council 

maintains a publication on their web site (www.awc.org) 

with listings of FSI data for domestic wood species. Only 

limited public data is available for the imported wood species. Further studies are needed to determine if the test environment of ASTM E 84 is applicable to potential fires in a 

covered bridge. It may be possible that Class B rated lumber 

may result in limited flame spread behavior for the typical 

fire scenarios that would result in minimal fire damage, with 

attendant low restoration costs. In any case, the test severity 

of ASTM E 84 usually ensures that the Class A rated lumber 

will provide the adequate passive protection from fire progression in realistic fire scenarios.

8.2 Pressure-Impregnated Fire-RetardantTreated Wood

Wood’s fire performance can be improved with fireretardant treatments. Commercial fire-retardant treatments 

for wood are intended to reduce the flame spread index as 

tested in accordance with ASTM E 84 (ASTM 2009a). In 

addition to slowing the spread of flames, the treatment will 

likely also delay ignition and reduce the heat release rate. 

The main treatment options are pressure impregnated treatment or surface application as a coating. Pressure treatments 

use pressure to impregnate the chemical solutions into wood 

members. This process is similar to the method used to impregnate chemical preservatives into wood members.

For building code purposes, specifications for FRTW go 

beyond the requirements for a Class A flame spread index 

of 25 or less. To qualify for the FRTW classification, the 

product must pass a modification of the standard E84 test in 

which the test is extended by 20 minutes to a total duration 

of 30 minutes. The test is often referred to as the 30-minute 

E 84 test and more recently has been described in ASTM E 

2768 (ASTM 2013a). Along with these added fire performance requirements, the requirements in the building codes 

for FRTW also require the treatment be a pressure-impregnated treatment and meet other performance requirements 

pertaining to hygroscopity and strength loss. Treated wood 

products that meet the requirements for FRTW can be used 

in applications where FRTW is a prescribed alternative material in the building code. The requirements for FRTW also 

include requirements for pre-weathering of the test specimen prior to the fire test when the application would subject 

the product to exterior weather exposures. Thus, specifications for FRTW need to specify whether the application is 

interior or exterior. In terms of the existing market, available 

FRTW products are limited to softwood lumber and softwood structural plywood. Available listings of species for 
Evaluating Fire-Damaged Components of Historic Covered Bridges

21

FRTW lumber are devoid of the hardwood species. FRTW 

products are required to have appropriate labels with information about the treatment.

The compositions of the commercial fire-retardant treatments are largely proprietary. Truax and others (1935) 

concluded that using different combinations of chemicals 

was the most effective way to provide fire-retardant treatments. Traditional treatments include the use of inorganic 

salts, such as boric acid, borax, diammonium phosphate, 

monoammonium phosphate, ammonium sulfate, and zinc 

chloride. Because the inorganic salts are subject to leaching, 

such treatments are limited to interior applications. Organic 

phosphate salts have also been used as a fire retardant on 

wood. The two common elements in FR treatments for 

wood are phosphorus and boron. While pressure-impregnated treatments are marketed as either a wood preservative or 

as a fire retardant, FR treatments incorporating boron will 

be likely to have improved resistance to decay and insect 

damage, as boron is also used as a wood preservative. Boron 

wood preservative treatments are likely to have little fire 

retardancy because the treatment levels for fire retardancy 

are much greater than those used for wood preservation.

As the result of high treatment levels, FR treatments are 

known to adversely impact the strength properties of the 

wood. In the structural design process, the reduced structural capacity is accounted for prescriptively through reduction factor for the allowable stresses. The reduction factors 

need to be obtained from the treater or the FR manufacturer. 

The specifications for FRTW also include test requirements 

intended to address potential for thermal degradation of the 

FRT product in roof or other applications involving elevated 

temperature exposures.

FR treatments for wood are not marketed to improve the fire 

resistance of a wood member as evaluated in the ASTM E 

119 test (ASTM 2010). The effect of the FR treatment on 

the charring rate is mixed. Many FR treatments lower the 

temperature required to initiate the thermal degradation of 

the wood. The result of this temperature shift is a reduction 

in the volatile gases and an increase in the residual char.

There are two published specifications for FRTW, namely 

NFPA 703: Standard for fire-retardant impregnated wood 

and fire-retardant coatings for building materials and the applicable sections of the AWPA standard (NFPA 2015). Manufacturers of FRTW as well as FR coatings may also have 

evaluation reports issued by the ICC Evaluation Service.

In applications where the building code provisions or other 

specifications are limited to just the requirements for the 

Class A or Class B flame spread index, the requirement for 

the FR treatment is limited to the FSI provision. In such 

cases, the use of a fire-retardant coating may be acceptable 

to the code official. FR composite products such as fiberboard and particleboard are likely to be marketed for just the 

Class A FSI provision. 

8.3 Fire Protective Coatings

For use in historic covered bridges, the FR coating is advantageous as a field application of the FR to the existing 

members. Pressure impregnation is obviously not an option 

for field application. Pressure impregnation would require 

removal of the wood for treatment or the use of new FRTW 

lumber. FR coatings can be sorted into two categories, intumescent and non-intumescent. Intumescent coatings are 

foam surfactants that expand when exposed to heat. When 

the foam expands, the volume increases while the density 

decreases and this provides a thermally insulative protective 

coating for the wood members. Specifications for FR coatings are incorporated in NFPA 703. FR coatings are marketed as meeting FSI of either Class A, B, or C. The classification will depend on the specifics of the wood substrate 

and the quantity of the application. As discussed previously, 

untreated lumber will typically meet either Class B or C 

depending on the species. Review of fire test documentation 

for a FR coating should take into consideration the likely 

classification of the uncoated substrate. As with any field 

application, quality control measures for the field application of the FR coating should be reviewed.

Similar to pressure impregnated FR treatments, FR coatings for wood for the most part are not marketed to stop the 

wood from charring. They are intended to reduce the spread 

of the initial flame. Fire protective coatings for structural 

steel are intended to improve their performance in the fire 

resistance test, ASTM E 119 (ASTM 2010). Research has 

shown that some fire protective coatings can in fact improve 

the fire resistance of wood members (White 1984, 1986; 

Richardson and Cornelissen 1984). These coatings are typically the intumescent coatings or thick coatings that can 

provide a degree of thermal insulation to the wood surface. 

More recently, the non-intumescent coatings tend to have an 

enhanced ability to absorb heat, such being hydrated with 

water, of which the brand “pyrotite” is one example. This 

FR coating is marketed as a 15-minute thermal barrier and 

as a 20-minute rating for use in assemblies. 

8.4 Panel Products

Gypsum board is the most common panel product used 

to improve the fire resistance (as measured by the ASTM 

E 119 test (ASTM 2010) of a wood member or assembly. 

Whereas regular gypsum board will provide improved fire 

resistance, it does not have fire resistance performance 

requirement. Fire rated gypsum board is known as Type 

X gypsum. Per its definition in ASTM C 1396/C 1396M 

(ASTM 2014), Type X designates gypsum board that provides not less than 1-hour fire resistance rating for boards 

5/8 in. thick, or 45 minutes fire-resistance rating for boards 

1/2 in. thick, as applied on each side of a typical wood 

frame wall assembly of 2 by 4 studs spaced at 16 in. on center. There is also a higher grade of fire-rated gypsum board 

that is often referred to as Type C. Gypsum board provides 
General Technical Report FPL–GTR–243

22

substantial protection because of its high content of bound 

water. In terms of performance, the differences between 

regular gypsum board and the different grades of fire-rated 

gypsum boards is their ability to maintain their own integrity after sustained fire exposure. During sustained fire exposure, cracks are formed in the gypsum board. 

8.5 Roof Coverings

The various types of roof coverings are rated for their relative fire performance. The series of fire tests that are used 

to classify roof coverings are specified in the standard 

ASTM E 108 (ASTM 2013b). The tests include those for 

flame spread as well as those intended to address fires being spread by burning brands or embers. The classifications 

for roof coverings are also Class A, B, and C, with Class 

A being the most restrictive. In the case of roof coverings 

and the ASTM E 108 test, untreated wood does not satisfy 

the requirements for Class C roof coverings. Fire-retardanttreated wood shingles are available that will satisfy Class C 

or Class B roof covering requirements. The roof shingles are 

typically western red cedar, a species that has some natural 

decay resistance. Wood roofs with Class A ratings are for the 

roof systems, which includes the specified underlayments 

(ASTM 2013b).

8.6 Security and Fire Alarm Options

The first means of protection against fire damage is prevention. Since historic covered bridges are often located in rural 

areas, fire-monitoring devices such as flame detectors can 

be used to alert officials of a possible threat of fire (Phares 

and others 2010). Faster response times minimize fire damage, thereby increasing the chances of successful post-fire 

rehabilitation (Fig. 11).

Minimizing fire damage to wood members is of the greatest 

importance to the rehabilitation, repair, and restoration of 

historic covered bridges. Because fire-retardant chemicals 

may reduce the structural capacity of wood members, the 

most effective way to minimize fire damage is to reduce 

burn time.

Historic covered timber bridges are often in remote locations. Fire on these structures will typically be reported after 

post-flashover has occurred or even after the decay period 

of a fire. Remote security and fire-detection monitoring system can be installed on all historic covered timber bridges 

to reduce response times; however, the cost effectiveness of 

these types of technologies should be considered.

A study carried out by Phares and others (2010) studied 

remote fire and vandalism monitoring systems using multiple fire detection devices and cameras on historic covered 

bridges. Iowa’s Cedar Covered Bridge was outfitted with an 

infrared camera, flame detector, and fiber optic sensors. The 

most expensive component of this system was the infrared 

camera, which cost $15,000. An infrared camera was used 

not only to detect trespassers after hours; it was also able to 

communicate with the other detection devices to minimize 

false fire alarms. Infrared cameras in this study were able to 

detect a pan-controlled fire in less than 10 seconds for each 

test and in less than 5 seconds for 75% of the tests. The pancontrolled fire is a contained fire in a 12-in.-diameter pan 

burning alcohol, because the standard gasoline fire was considered too volatile for use on a historic covered bridge.

Flame detectors also proved to be a relatively reliable fire 

detector, as they were able to detect a pan-controlled fire in 

7 seconds at a distance of 15 ft; however, they were unable 

to detect lit butane lighter beyond 24 in. Even though the 

detection of small fires such as a lighter may be a limitation, 

the detection times associated with this technology proved 

to be successful. By installing fiber optic sensors near the 

timber bridge deck, they were also able to detect the presence of a fire when the fire was within a few feet of the 

sensor. This may be beneficial, as most fires are lit on the 

bridge deck alongside an interior wall of the bridge.

Of the three devices tested, the flame detectors proved to be 

the method of choice for detecting historic covered timber 

bridge fires because of their quick detection time for fire, 

ease of installation, reliability, and their relatively low cost.

8.7 Fire Extinguishment Options

The most popular forms of fire extinguishment include suppressing agents in foam or powder form, carbon dioxide, 

and water. All of these forms can be administered in different manners to increase the firefighting effectiveness of the 

Figure 11—Examples of fire deterrent and notification techniques used on the Red Covered Bridge, circa 2012.
Evaluating Fire-Damaged Components of Historic Covered Bridges

23

media being used. Examples include but certainly aren’t 

limited to automated systems, water additives, fire hydrants, 

and portable fire extinguishers.

Depending on the locations of historic covered bridges, the 

standard fire extinguishment may very well be through automated systems. Pipe and nozzle systems have long been 

used alongside water. Environmental impact is minimized 

because water, while readily available in most instances, 

is also inert. Water application systems must be properly 

maintained to avoid unnecessary exposure-related issues 

leading to the degradation of wood components. Water-mist 

applications can be differentiated into one of three types of 

systems: total, local, and zoned.

With respect to water additives, the advantages of a multicomposition (MC) additive to water-dispersing systems are 

well documented. The MC additives are made from mixing 

five different components. Of the additives, a fluorocarbon 

surfactant is used to form a thin film layer on members that 

are in line with the spray. A viscosity modifier can be used 

to improve the mobility and runoff of the mist spray; this 

allows for a quicker dispersion of the water and additive. An 

organic metallic compound can also be added to improve 

the fire extinguishing capabilities of the water by producing 

active radicals. Carbamide is used to absorb energy from 

the flame. It generates inert gases as a byproduct. The final 

compound, N,N-dimethylformamide, serves as an antifreeze. Collectively, these chemicals, when added to a water 

mist spray, allow for a more effective way of extinguishing 

a remote bridge fire before irreparable damage to the bridge 

occurs (Zhou and Guangxuan 2006).

Should remote systems prove to be prohibitive, fire hydrants 

have also been used to extinguish bridge fires. Water must 

be supplied to the fire hydrant, which may or may not be a 

practical solution for all bridges in question. Similarly, portable fire extinguishers require an operator. Fire detection 

and response are not addressed through approaches such 

as this. Concerns of remoteness and readiness remain. As 

standalones, techniques like this could be construed as a last 

resort option for fighting, but have real advantages when 

incorporated into a larger extinguishment plan.

9.0 Concluding Remarks and 

Summary

Arson continues to claim many historic covered bridges. 

Site-specific post-fire evaluations of the structural integrity 

of a bridge are often necessary in a fire’s aftermath. Decisions must be made on whether individual wood components can be rehabilitated, reconstructed, or replaced. This 

project on evaluating fire-damaged components of historic 

covered bridges includes post-fire assessment guidelines for 

evaluations, existing methodologies and techniques, and the 

customized application of a nondestructive technique, all of 

which can be employed during site-specific post-fire evaluations of historic covered bridges. More specifically,

• The existing approaches to post-fire evaluations were 

surveyed based on their relevance to historic covered 

bridges. A focal point included the rehabilitation, reconstruction, and replacement of individual fire-damaged 

flexural members.

• A nondestructive technique for directly determining the 

residual capacity of individual fire-damaged glued-laminated beams that are expected to remain in service was 

also produced.

10.0 Federal Program for Covered 

Bridges

Several important federal agencies participate in the preservation and registry of historic covered bridges. The largest 

database is the National Register of Historic Places. Owners 

and funding agencies alike can apply to register a bridge, an 

important step in the preservation and repair of such bridges.

The Historic American Engineering Record (HAER), part 

of the National Park Service, develops well documented 

records of historic structures including covered bridges. 

The HAER collection on covered bridges houses more than 

3,500 sheets of measured and interpretive drawings, 72,000 

photographs, and 61,000 data pages on over 7,000 sites, 

structures, and objects (Pierce and others 2005). All of these 

forms of data can be found at the U.S. Library of Congress, 

or the National Digital Library.

The U.S. Secretary of Interior’s Standards for Historic Preservation has provided the Standards and Guidelines for Archeology and Historic Preservation (Federal Register 1983). 

These guidelines provide technical advice on preservation 

activities and methodologies including, but not limited to, 

rehabilitation, reconstruction, and restoration.

The FPL actively tests and evaluates different aspects of 

wood bridge design and fire-damage evaluation. The tests 

conducted by FPL allow for technological advances that 

would not otherwise be possible. The works of this agency 

are far reaching. For more extensive information on these 

and other agencies and/or organizations, refer to Chapter 18 

of the Covered Bridge Manual (Pierce and others 2005).

The Covered Bridge Manual (Pierce and others 2005) provides excellent comprehensive support to those involved 

with assessing, maintaining, and rehabilitating covered 

bridges. While this resource states that site-specific post-fire 

evaluations are necessary in a fire’s aftermath, additional 

guidance is absent. In particular, there is no direction on 

how to decide whether individual wood components can be 

rehabilitated, reconstructed, or replaced. Recognizing the 

need for additional information on this topic, this report is 

thought to complement and enhance the Covered Bridge 
General Technical Report FPL–GTR–243

24

Manual current coverage of post-fire evaluations. It is hoped 

to be worthy of being included in the Covered Bridge Manual or similar future publications of the National Historic 

Covered Bridge Preservation Program.

In terms of historic preservation, engineers are to comply 

with the Historic Sites Act of 1935, Section 106 of the 

National Historic Preservation Act of 1966 and its amendments. The Federal Highway Administration’s National 

Historic Covered Bridge Preservation Program started with 

its highway legislation the Transportation Equity Act for the 

21st Century (TEA 21) (FHA 1999). It was subsequently 

renewed under the next Highway Bill the Safe, Accountable, 

Flexible, Efficient Transportation Equity Act: A Legacy for 

Users (SAFETEA-LU) (FHA 2009–2012), which funded 

rehabilitation, restoration, and preservation of covered 

bridges listed or eligible for listing on the National Register 

of Historic Places, and research and education. Information 

pertaining to the National Register for Historic Places application process can be found on the United States Department of Transportation website under their National Historic 

Covered Bridge Preservation Program (http://www.fhwa.

dot.gov/bridge/covered.cfm). Procedures that relate to historic covered bridge improvements such as strengthening 

also need to conform to the Secretary of Interior’s guidelines for preservation of historic structures.

The TEA 21 and SAFETEA-LU transportation bills provided funding for the repair and preservation of historic covered bridges, and for research, but did not provide funds for 

the full reconstruction of a historic bridge if it cannot be repaired (Wacker and Duwadi 2008). This program, however, 

expired with the expiration of SAFETEA-LU. With the loss 

of this approximately $10 million per year program, repairing and rehabilitating historic covered bridges becomes an 

even more difficult task. Preservationist groups could still 

apply for historic covered bridge grants, but it would be necessary to compete directly with other bridge and highway 

projects for funds. Therefore, the rehabilitation and reinforcing of historic covered bridges damaged by fire is extremely 

important to saving these historic landmarks from destruction after a fire event.

For timber bridges listed under the National Historic Registry (www.nps.gov/nr/), post-fire repair should aim to satisfy 

strength requirements from a structural, aesthetic, and historical perspective. Whenever possible, the strengthening 

methods should conform to the Secretary of Interior’s guidelines for preservation of historic structures (Federal Register 

1983).

The Vermont Historic Bridge Program (2014) has developed 

a progressive list of rehabilitation techniques. When such 

techniques are warranted, their adoption helps ensure that 

covered timber bridges accommodate historically-accurate 

engineering procedures that utilize similar materials. The 

Vermont Historic Covered Bridge Program mentions that if 

this is not feasible, member augmentation may be used after 

other options have been evaluated and are otherwise ruled 

out. In this case, modern materials like glass-fiber reinforced plastics (GFRP) exhibit extremely low profiles once 

installed and have been shown to increase the load-carrying 

capacity (Kim and Harries 2010; Gentile and others 1999).

References

Anderson, C.A.; Loferski, J.R.; Woeste, P.E. 2003. Detecting 

early wood decay. The Inspector’s Journal, Online. Reprint 

accessed on 08/25/14 at http://www.inspectorsjournal.com/

forum/topic.asp?TOPIC_ID=2101&SearchTerms=pick+test

American Lumber Standard Committee, Inc. 2015. American Softwood Lumber Standard Rules. Germantown, Maryland: American Lumber Standard Committee, Incorporated. 

http://www.alsc.org/greenbook%20collection/LumberProgram_facsimile.pdf

American Wood Council. 2012. National Design Specification for wood construction with commentary. ASD/LRFD. 

Washington D.C.: American Wood Council.

American Wood Council. 2014. Calculating the fire resistance of exposed wood members: Technical Report 10. 

Washington D.C.: American Wood Council. 51 p.

APA. 1989. Fire retardant treated plywood roof sheathing: In-situ testing. Tacoma, WA: American Plywood 

Association.

APA. 2009. Calculating fire resistance of glulam beams and 

columns. EWS Y245B. Tacoma, Washington: American 

Plywood Association – The Engineered Wood Association. 

http://www.drjlumber.com/wp-content/uploads/2014/07/

Calculating-Fire-Resistance.pdf

ASTM. 1992. Standard test methods for use and calibration 

of hand-held moisture meters. Designation D4444-92. In: 

Annual book of ASTM standards. West Conshohocken, PA: 

American Society for Testing and Materials International.

ASTM. 1994. Standard test methods for small clear specimens of timber. ASTM D143-94. West Conshohocken, PA: 

American Society for Testing and Materials International.

ASTM. 1999. Standard guide for moisture conditioning of 

wood and wood-based materials. Designation D4933-99. In: 

Annual book of ASTM standards. West Conshohocken, PA: 

American Society for Testing and Materials International.

ASTM. 2000. Standard test method for structural panels in 

flexure. Designation D3043 In: Annual book of ASTM standards. West Conshohocken, PA: American Society for Testing and Materials International.

ASTM. 2009a. Standard test method for surface burning characteristics of building materials. Designation 

E84-09. In: Annual book of ASTM standards. West Conshohocken, PA: American Society for Testing and Materials 

International.
Evaluating Fire-Damaged Components of Historic Covered Bridges

25

ASTM. 2009b. Standard test methods of static tests of lumber in structural sizes. Designation D198-09. In: Annual 

book of ASTM standards. West Conshohocken, PA: American Society for Testing and Materials International.

ASTM. 2010. Standard test methods for fire tests of building construction and materials. Designation E119-10a. In: 

Annual book of ASTM standards. West Conshohocken, PA: 

American Society for Testing and Materials International.

ASTM. 2013a. Standard test method for extended duration 

surface burning characteristics of building materials (30 

min tunnel test). Designation E2768-11. In: Annual book of 

ASTM standards. West Conshohocken, PA: American Society for Testing and Materials International.

ASTM. 2013b. Standard test methods for fire tests of roof 

covering. Designation E108-11. In: Annual book of ASTM 

standards. West Conshohocken, PA: American Society for 

Testing and Materials International.

ASTM. 2014. Standard specification for gypsum board. 

Designation C1396/C1396M-14a. In: Annual book of 

ASTM standards.West Conshohocken, PA: American Society for Testing and Materials International.

Barbaresi, T.; Cozzani, V.; Di Padova, A.; Tallone, F.; 

Tugnoli, A. 2012. Mitigation of fire damage and escalation 

by fireproofing: A risk-based strategy. Reliability Engineering and System Safety. 105: 25.

Babrauskas, V. 2001. Ignition of wood: a review of the state 

of the art. London: Interflam, Interscience Communications, 

Ltd.: 71–88.

Barletta, M. 2005. Application of instrumented micro-indentations to ‘in situ’ mechanical characterization of wooden 

structures: Part I—Analysis of highly selected and decayed 

pinewood samples. Measurement. 39(4): 274–284.

Borri, A.; Corradi, M. 2011. Strengthening of timber beams with high strength steel cords. Composites 

Part B: Engineering. 42(6): 1480–1491. doi:10.1016/j.

compositesb.2011.04.051

Branco, J.; Piazza, M.; Cruz, P. 2009. Structural analysis of 

two King-post timber trusses: Non-destructive evaluation 

and load-carrying tests. Construction and Building Materials. 24(3): 371–383.

Brashaw, B.K.; Bucur, V.; Divos, F.; Gonҫalves, R.; Lu, J.; 

Meder, R.; Pellerin, R.; Potter, S.; Ross, R.; Wang, X.; Yin, 

Y. 2009. Nondestructive testing and evaluation of wood: A 

worldwide research update. Forest Products Journal. 59(3): 

7–14. 

Brashaw, B.K.; Wang, X.; Ross, R.J.; Pellerin, R.F. 2004. 

Relationship between stress wave velocities of green and 

dry veneer. Forest Products Journal. 54(6): 85–89. 

Browne, F.L. 1958. Theories of the combustion of wood 

and its control. FPL Report 2136. Madison, WI: U.S. 

Department of Agriculture, Forest Service, Forest Products 

Laboratory.

Buchanan, A.H. 2001. Structural design for fire safety. New 

York: John Wiley & Sons, Ltd.

Buchanan, A.; Barber, D. 1994. Fire resistance of epoxied 

steel rods in glulam timber. In: Proceedings, Volume 1, 

Pacific Timber Engineering Conference; 1994 July 11–15; 

Gold Coast, Australia. Fortitude Valley MAC, Queensland, 

Australia: Timber Research and Development Advisory 

Council: 590–598.

Custódio, J.; Broughton, J.; Cruz, H. 2009. A review of factors influencing the durability of structural bonded timber 

joints. International Journal of Adhesion and Adhesives. 

29(2): 173–185. doi:10.1016/j.ijadhadh.2008.03.002

Dietenberger, M.A. 2004. Ignitability of materials in transitional heating regimes. In: Wood and fire safety. Proceedings, 5th international scientific conference, April 18–22, 

2004. Slovak Republic, Svolen, Slovakia: Faculty of Wood 

Sciences and Technology, Technical University of Zvolen: 

31–41.

DOT. 2007. Strengthening historic bridges to carry modern traffic. FHWA-HRT-07-041. McLean, VA: United 

States Department of Transportation, Federal Highway 

Administration.

Emerson, R.N.; Pollock, D.G.; Kainz, J.A.; Fridley, K.J.; 

McLean, D.L.; Ross, R.J. 1998. Nondestructive evaluation 

techniques for timber bridges. In: Proceedings of the 5th 

world conference on timber engineering, Swiss Federal Institute of Technology, Lausanne, Switzerland: 670–677.

Esteves B.; Pereira, H. 2008. Quality assessment of heattreated wood by NIR spectroscopy. Holz als Roh- und 

Werkstoff. 66(5): 323–332. 

Federal Register. 1983. Thursday, September 29, 1983, 

Part IV, Department of the Interior, National Park Service, 

Archeology and Historic Preservation: Secretary of the Interior’s Standards and Guidelines. 48(190): 44738–44739.

Felicetti, Roberto. 2006. The drilling resistance test for the 

assessment of fire damaged concrete. Cement & Concrete 

Composites. 28(2006): 321–329.

FHA. 1999. Public Law 105–178, as amended by Title IX of 

Public Law 105–206. Transportation Equity Act for the 21st 

century. TEA 21. Washington, D.C.: U.S. Federal Highway 

Administration. https://www.fhwa.dot.gov/tea21/h2400idx.

htm#sec1

FHA. 2009–2012. Safe, Accountable, Flexible, Efficient 

Transportation Equity Act: A legacy for users (SAFETEALU) Extension Acts. Washington, D.C.: U.S. Federal Highway Administration. https://www.fhwa.dot.gov/safetealu/

legis.htm
General Technical Report FPL–GTR–243

26

Firmanti, A.; Backtiar E.T.; Surjokusumo, S.; Komatsu, K.; 

Kawai, S. 2004. Mechanical stress grading of tropical timbers without regard to species. Journal of Wood Science. 51: 

339–347.

Fox, S.P. 1974. Residual strength of fire-damaged laminated 

beams. Forest Products Journal. 24(1): 48–50.

Freas, A.D. 1982. Evaluation, maintenance and upgrading 

of wood structures: a guide and commentary. American Society of Civil Engineers, Committee on Wood, Subcommittee on Evaluation, Maintenance and Upgrading of Timber 

Structures. New York: American Society of Civil Engineers. 

428 p.

Fuller, J.J.; Leichti, R.J.; White, R.H. 1992. Temperature 

distribution in a nailed gypsum-stud joint exposed to fire. 

Fire and Materials. 16: 95–99.

Garab, J.; Toth, A.; Szalai, J.; Bejo, L.; Divos, F. 2010. Evaluating glued laminated beams using a nondestructive testing 

technique. Transactions of FAMENA. 34(4): 33–46.

Garcia, R.A.; Carvalho, A.M.; Latorraca, V.F.; Matos, 

J.L.M.; Santos, W.A.; Silva, R.F.M. 2012. Nondestructive evaluation of heat-treated Eucalyptus grandis Hill ex 

Maiden wood using stress wave method. Wood and Science 

Technology. 46:41–52.

Gentile C.; Svecova, D.; Saltzberg W.; Rizkalla, S. 1999. 

Strengthening of timber bridges. World Wise ’99, Winnipeg, 

Canada, December 6–8, 1999. 6 p.

Glos, P.; Henrici, D. 1991. Bending strength and MOE of 

structural timber (Picea abies) at temperatures up to 150°C. 

Holz als Roh-und Werkstoff. 49(1991): 417–422. Translated 

by Barbara J. Marshment, May 1992.

Green, D.W.; Evans, J.W. 2008. The immediate effect of 

temperature on the modulus of elasticity of green and dry 

lumber. Wood and Fiber Science. 40(3): 374–383.

Green, D.W.; Winandy, J.E.; Kretschmann, D.E. 1999. Mechanical properties of wood–wood as an engineering material. General Technical Report FPL–GTR–113. Madison, 

WI: U.S. Department of Agriculture, Forest Service, Forest 

Products Laboratory.

Griggs, Frank. 2014. Union Bridge: First bridge across the 

Hudson River, 1804. Structure Magazine. February 2014.

ISIS Canada. 2000. Intelligent Sensing for Innovative Solutions (ISIS) Canada. Solution-oriented research. Annual 

Report (2000–2001). ISIS Canada Administrative Centre, 

The University of Manitoba, Winnipeg, Manitoba, Canada. 

http://www.isiscanada.com/archived-isiscanada/annual_report/ARE_01.pdf

James, W.L. 1988. Electric moisture meters for wood. General Technical Report FPL–GTR–6. Madison, WI: U.S. 

Department of Agriculture, Forest Service, Forest Products 

Laboratory. 17 p.

Janssens, M.L.; White R.H. 1994. Short communication: 

temperature profiles in wood members exposed to fire. Fire 

and Materials. 18: 263–265.

Kim, Y.J.; Harries, K.A. 2010. Modeling of timber beams 

strengthened with various CFRP composites. Engineering Structures. 32(10): 3225-3234. doi:10.1016/j.

engstruct.2010.06.011

King, M. 2002. NIDR guidelines for fire and smoke damage 

repair. Millerville, MD: The National Institute of Disaster 

Restoration.

Klinkhachorn, P.; Halabe, U.B.; GangaRao, H. 1999. Portable ultrasonic device to monitor timber properties: Assessment of charred timber members. In: Proceedings of the 

28th International Conference on Fire Safety, Columbus, 

OH: 89–93.

Kretschmann, D.E. 2010. Mechanical properties of wood. 

In: Wood handbook, wood as an engineering material. General Technical Report FPL–GTR–190. Madison, WI: U.S. 

Department of Agriculture, Forest Service, Forest Products 

Laboratory.

Kukay, B.M.; Barr, P.J.; Friel, L.; Coster, D.C.; Halling, 

M.W. 2008. Post-fire assessments, methodology, and equations for directly determining wood’s residual flexural properties. Forest Products Journal. 58(1–2): 40–46.

Kukay, B.M.; Todd, C. 2009. Determining wood’s residual 

flexural properties using nondestructive testing. Forest Products Journal, Wood Design Focus. 19(2): 8–12.

Kukay, B.; White, R.; Todd, C.; Jahn, T. 2013. Evaluating 

fire-damaged components of timber bridges and structures. 

Structures Congress 2013, Fire Risks and Impacts to Bridges Track. Pittsburgh, PA: 758–767.

Lebow, S.T.; Anthony, R.W. 2012. Guide for use of wood 

preservatives in historic structures General Technical Report 

FPL–GTR–217. Madison, WI: U.S. Department of Agriculture, Forest Service, Forest Products Laboratory. 59 p.

Lebow, S.; Kirker, G.; White, R.; Amburgey, T.; Barnes, 

H.M.; Sanders, M.; Morrell, J. 2012. Guide for in-place 

treatment of wood in historic covered and modern bridges. 

General Technical Report FPL–GTR–205. Madison, WI: 

U.S. Department of Agriculture, Forest Service, Forest 

Products Laboratory. 43 p.

Lebow, S.T.; Winandy, J. E. 1999. Effect of fire-retardant 

treatment on plywood pH and the relationship of pH to 

strength properties. Wood Science and Technology. 33(4): 

285–298.

LeVan, S.L.; Winandy, J.E. 1990. Effects of fire retardant 

chemical on the bending properties of wood at elevated temperatures. Research Paper FPL–RP–498. Madison, WI: U.S. 

Department of Agriculture, Forest Service, Forest Products 

Laboratory.
Evaluating Fire-Damaged Components of Historic Covered Bridges

27

NAHB. 1990. Home builders guide to fire retardant treated 

plywood. Evaluation, testing and replacement. Upper Marlboro, MD: National Association of Home Builders, National 

Research Center. 49 p. http://www.awc.org/helpoutreach/

faq/FAQfiles/FRT-Publications-listandref.html

NFPA. 2015. Standard for fire retardant-treated wood and 

fire-retardant coatings for building materials. Standard 703. 

Quincy, MA: National Fire Protection Association.

NFPA. 2013. Guide for fire and explosion investigations. 

NFPA 921. Quincy, MA: National Fire Protection Association. 305 p.

Nicholas, D.D.; Crawford, D. 2003. Concepts in the development of new accelerated test methods for wood decay. In: 

ACS Symposium Series. 845: 288–312. Washington, DC: 

American Chemical Society.

Niemz, P.; Mannes, D. 2012. Non-destructive testing of 

wood and wood-based materials. Journal of Cultural Heritage. 13(3): S26–S34, Supplement.

Nuopponen, M.; Wikberg, H.; Vuorinen, T.; Maunu, S.L.; 

Jämsä, S.; Viitaniemi, P. 2004. Heat-treated softwood exposed to weathering. Journal of Applied Polymer Science. 

91: 2128–2134. 

Pellerin, R.F.; Lavinder, J.A.; Ross, R.J.; Falk, R.H. 1996. 

NDE of timber bridges. In: Proceedings, SPIE, Nondestructive Evaluation of Materials and Composites, November 15, 

1996. Volume 2944: 275–284. doi:10.1117/12.259068

Phares, Brent M.; LaViolette, Michael D.; Wipf, Terry J.; 

Ritter, Michael A. 2010. Remote monitoring of historic covered timber bridges for the prevention of arson and vandalism. General Technical Report FPL–GTR–191. Madison, 

WI: U.S. Department of Agriculture, Forest Service, Forest 

Products Laboratory. 27 p.

Phares, B.; Wipf, T.; Sievers, R.; Hosteng, T. 2013. Covered 

bridge security manual. General Technical Report FPL–

GTR–223. Madison, WI: U.S. Department of Agriculture, 

Forest Service, Forest Products Laboratory. 34 p.

Pierce, P.C.; Brungraber, R.L.; Lichtenstein, A.; Sabol, 

S. 2005. Covered bridge manual. FHWA-HRT-04-098. 

McLean, Virginia: Federal Highway Administration. http://

www.fhwa.dot.gov/publications/research/infrastructure/

structures/04098/index.cfm. 

Raftery, G.M.; Harte, A.M.; Rodd, P.D. 2009. Bond quality 

at the FRP wood interface using wood-laminating adhesives.

International Journal of Adhesion and Adhesives. 29(2): 

101–110. doi:10.1016/j.ijadhadh.2008.01.006

Reinprecht, L.; Panek, M. 2012. Ultrasonic technique for 

evaluation of biodefects in wood: Part 1 – influence of the 

position, extent, and degree of artificial rots. International 

Wood Products Journal. 3(2): 107–115.

Richardson, L.R.; Cornelissen, A.A. 1984. Fire-resistant 

coatings for roof/ceiling deck timbers. Fire and Materials. 

11: 191–194.

Ritter, M.A. 1992. Maintenance and rehabilitation of timber 

bridges. In: Timber bridges: design, construction, inspection, and maintenance. Chapter 14. Washington, DC: U.S. 

Department of Agriculture, Forest Service.

Rodgers, P.; Otter, D.E.; Wipf, T.; Ritter, M.; Wood, D. 

1997. Effectiveness of glued laminated components in 

strengthening timber bridges. Technical Document 

TD 97-026. Washington D.C.: Association of American 

Railroads.

Ross, R.J.; Brashaw, B.K.; Wang, X.; White, R.H.; Pellerin, 

R.F. 2005a. Post-fire assessment of structural wood members. In: Wood and timber condition assessment manual. 

Chapter 4. Madison, WI: Forest Products Society: 29–46. 

Ross, R.J.; Wang, X.; Brashaw, B.K. 2005b. Detecting 

decay in wood components. In: Inspection and monitoring 

techniques for bridges and civil structures. Cambridge, England: Woodhead Publishing Limited: 100–114.

Ross, R.J.; Cooper, J.; Wang, Z. 1991a. In-place evaluation 

of fire retardant treated wood products. In: Proceedings, 

Eighth International Nondestructive Testing of Wood Symposium, Pullman, WA: 247–252.

Ross, R.J.; Pellerin, R.F. 1994. Nondestructive testing for 

assessing wood members in structures: A review. General 

Technical Report FPL–GTR–70. Madison, WI: U.S. Department of Agriculture, Forest Service, Forest Products 

Laboratory.

Ross, R.; Pellerin, R.F.; Volny, N.; Salsig, W.W.; Falk, R.H. 

2000. Stress wave timing nondestructive evaluation tools for 

inspecting historic structures: A guide for use and interpretation. General Technical Report FPL–GTR–119. Madison, 

WI: U.S. Department of Agriculture, Forest Service, Forest 

Products Laboratory.

Rowell, R.M.; Dietenberger, M.A. 2013. Thermal properties, combustion, and fire retardancy of wood. In: Rowell, 

R., ed. Handbook of wood chemistry and wood composites. 

Second edition. New York: CRC Press. pp. 127–149. 

Chapter 6.

Schaffer, E. 1977. State of structural timber fire endurance. 

Wood and Fiber. 9(2): 145–170.

Schaffer, E.L. 1982a. Temperature. In: Evaluation, maintenance and upgrading of wood structures: a guide and commentary. New York: American Society of Civil Engineers: 

69–92. Chapter 2, Section 2.3. 

Schaffer, E.L. 1982b. Fire. In: Evaluation, maintenance and 

upgrading of wood structures: a guide and commentary. 

New York: American Society of Civil Engineers: 120–126. 

Chapter 2, Section 2.7.
General Technical Report FPL–GTR–243

28

Schaffer, E.L. 1984. Structural fire design: wood. Research 

Paper FPL 450. Madison, WI: U.S. Department of Agriculture, Forest Service, Forest Products Laboratory. 16 p.

Schroeder, R.A. 1999. Post-fire analysis of construction 

materials. Berkeley, CA: University of California Berkeley. 

Ph.D. dissertation. 61–03, 1–1–275. 

So, C.L.; Via, B.K.; Groom, L.H.; Schimleck, L.R.; Shupe, 

T.F.; Kelley, S.S.; Rials, T.G. 2004. Near infrared spectroscopy in the forest products industry. Forest Products Journal. 54(3): 6–16.

Truax, T.R.; Harrison, C.A.; Baechler, R.H. 1935. Experiments in fireproofing wood—fifth progress report. In: Proceedings of the American Wood Preservers’ Association: 

231–245.

Valipour, H.R.; Crews, K. 2011. Efficient finite element 

modeling of timber beams strengthened with bonded fibre 

reinforced polymers. Construction and Building Materials. 

25(8): 3291–3300. doi:10.1016/j.conbuildmat.2011.03.017.

Vermont Historic Bridge Program. 2014. A Vermont Government Website. Copyright 2014. State of Vermont. http://

historicbridges.vermont.gov/

Wacker, J.P.; Duwadi, S.R. 2008. Covered bridges in the 

United States and the preservation program. Federal Highway Administration. In: Proceedings of the 10th world conference on timber engineering. Miyazak, Japan; 2008 June 

2–5. 8 p.

Wang, N.; Fu, Y.; Liu Y.; Yu, H.; Liu, Y. 2014. Synthesis of 

aluminum hydroxide thin coating and its influence on the 

thermochemical and fire-resistant properties of wood. Holzforschung. 68(7): 781–789.

Wang, X.; Wacker, J.P.; Ross, R.J.; Brashaw, B.K. 2005. 

Nondestructive assessment of timber bridges using a vibration-based method. In: Proceedings of the 2005 structures 

congress. New York: The Structural Engineering Institute of 

the American Society of Civil Engineers.

Watanabe, K.; Hart, J.F.; Mansfield, S.D.; Avramidis, S. 

2010. The future of quality control for wood and wood 

products. In: Near-infrared technology applications for quality control in wood processing. Proceedings, The Final Conference of COST Action E53, 4–7, May 2010, Edinburgh.

White, R.H. 1984. Use of coating to improve fire resistance 

of wood. In: Lieff, M.; Stumpf, F.M., eds. Fire resistive 

coatings: the need for standards. ASTM STP 826. Philadelphia: American Society for Testing and Materials: 24–39.

White, R.H. 1986. An empirical model for predicting performance of fire-resistive coatings in wood construction. Journal of Testing and Evaluation. 14(2): 97–108. 

White, R.H. 2004. Post-fire assessment of structural wood 

members. In: Wood and timber condition assessment 

manual, first edition. Madison, WI: Forest Products Society: 

29–46. Chapter 4.

White, R.H.; Cramer, S.M.; Shrestha, D.K. 1993. Fire endurance model for a metal-plate-connect wood truss. Research Paper FPL–RP–522. Madison, WI: U.S. Department 

of Agriculture, Forest Service, Forest Products Laboratory. 

12 p.

White, R.H.; Dietenberger, M.A. 2001.Wood products: thermal degradation and fire. In: Encyclopedia of materials: science and technology. Elsevier Science, Ltd.: 9712–9716.

White, R.H.; Dietenberger, M.A. 2010. Fire safety of wood 

construction. In: Wood handbook—wood as an engineering 

material. General Technical Report FPL–GTR–190. Madison, WI: U.S. Department of Agriculture, Forest Service, 

Forest Products Laboratory: 18-1–18-22.

White, R.H.; Kukay, B.; Wacker, J. 2013a. Options for NDE 

assessment of heat and fire-damaged wood. 18th International Nondestructive Testing and Evaluation of Wood Symposium, September 2013, Madison, Wisconsin.

White, R.H.; Kukay, B.; Wacker, J.; Sannon, J. 2013b. Options for evaluating fire-damaged components of historic 

covered bridges. Second National Covered Bridge Conference. Dayton, OH.

White, R.H.; Nordheim, E.V. 1992. Charring rate of wood 

for ASTM E 119 exposure. Fire Technology. 28(1): 5–30. 

White, R.H.; Ross, R.J. (Eds.). 2014. Wood and timber condition assessment manual: second edition. General Technical Report FPL–GTR–234. Madison, WI: U.S. Department 

of Agriculture, Forest Service, Forest Products Laboratory. 

93 p.

White, R.H.; Woeste, F. 2013. Post-fire analysis of solidsawn heavy beams. Structure Magazine. November 2013.

Williamson, T.G. 1982a. Rehabilitation of fire-damaged 

members.In: Evaluation, maintenance and upgrading of 

wood structures: a guide and commentary. New York: 

American Society of Civil Engineers: 172–178. Chapter 4, 

Section 4.5.4.

Williamson, T.G. 1982b. Rehabilitation of fire-damaged 

timber—the Filene Center. In: Evaluation, maintenance and 

upgrading of wood structures: a guide and commentary. 

New York: American Society of Civil Engineers: 417–422. 

Chapter 8, Section 8.8.

Winandy, J.E. 2001. Thermal degradation of fire-retardanttreated wood: predicting residual service-life. Forest Products Journal. 51(2): 47–54.

Winandy, J.E.; Ross, R.J.; LeVan, S.L. 1991a. Fire retardant 

treated wood: Research at U.S. Department of Agriculture, 

Forest Service, Forest Products Laboratory. Proceedings, International Timber Engineering Conference, London, 1991 

Sept. 1–5. p. 4.69–4.74. 
Evaluating Fire-Damaged Components of Historic Covered Bridges

29

Winandy, J.E.; LeVan, S.L.; Ross, R.J.; Hoffman, S.P.; McIntyre, C.R. 1991b. Development and evaluation of a test 

protocol to compare the effects of extended exposure to 

elevated temperature on untreated and fire-retardant-treated 

plywood. Research Paper FPL–RP–501. Madison, WI: U.S. 

Department of Agriculture, Forest Service, Forest Products 

Laboratory.

Winandy, J.E.; Lebow, P.K.; Nelson, W. 1998. Predicting 

bending strength of fire-retardant-treated plywood from 

screw-withdrawal tests. Research Paper FPL–RP–568. Madison, WI: U.S. Department of Agriculture, Forest Service, 

Forest Products Laboratory. 20 p.

Yahyaei-Moayyed, M.; Taheri, F. 2011. Experimental and 

computational investigations into creep response of AFRP 

reinforced timber beams. Composite Structures. 93(2): 616–

628. doi:10.1016/j.compstruct.2010.08.017

Zicherman, J.B.; Williamson, R.B. 1981. Microstructure of 

wood char. Wood Science and Technology. 15: 237–249.

Zickler, G.A.; Schoberl, T.; Paris, O. 2006. Mechanical 

properties of pyrolysed wood: a nanoindentation study. Philosophical Magazine A: Physics of condensed matter: structure, defects and mechanical properties. 86(10): 1373–1386.

Zhou, X.; Guangxuan, L.; Bo, C. 2006. Improvement of 

water mist’s fire-extinguishing efficiency with MC additive. 

Fire Safety Journal. 41: 39–45.

Additional Sources 

American Forest & Paper Association. 2001. The allowable 

stress design (ASD) manual for engineered wood construction. Washington DC: American Forest & Paper Association.

American Wood Council. 2014. Design for code acceptance. 

http://www.awc.org/Publications/dca/dca1/DCA1.pdf 

ASTM. 2010. Practice for sampling and data-analysis for 

structural wood and wood-based products. Designation 

D2915-10. In: Annual book of ASTM standards. West Conshohocken, PA: American Society for Testing and Materials 

International.

ASTM. 2009. Standard test method for evaluating the 

flexural properties of fire-retardant treated softwood 

plywood exposed to elevated temperatures. Designation 

D5516-09. In: Annual book of ASTM standards. West Conshohocken, PA: American Society for Testing and Materials 

International.

ASTM. 2010. Standard test method for evaluating the effects of fire-retardant treatments and elevated temperatures 

on strength properties of fire-retardant treated lumber. Designation D5664-10. In: Annual book of ASTM standards. 

West Conshohocken, PA: American Society for Testing and 

Materials International.

ASTM. 2013. Standard test methods for standardization and 

calibration of in-line dry lumber moisture meters. Designation D6782-13. In: Annual book of ASTM standards. West 

Conshohocken, PA: American Society for Testing and Materials International.

AWPA. 2014. AWPA book of standards. AWPA Technical 

Committee. American Wood Protection Association.

White, R.H. 2008. Analytical methods for determining fire 

resistance of timber members. In: SFPE handbook of fire 

protection engineering, fourth edition. Quincy, MA: National Fire Protection Association: 346–366.
General Technical Report FPL–GTR–243

30

A.1 Technical Basis for Technique

The goal of this research is to adapt a direct test to estimate the residual flexural properties of individual framing 

members for post-fire evaluations. The concentrated proof 

loading tests are typically conducted in the field. The proof 

loading technique summarized below is similar to that 

which is mentioned in the NAHB FRT evaluation guide, 

created by the American Plywood Association for untreated 

plywood panels (NAHB 1990). The adaptation incorporates 

partial-composite action with an aluminum stock and a 

cycle of increasing load tests that are applied gradually. The 

evaluations of the methodology were based on eight, gluedlaminated beams of which four were fire-damaged. The 

fire-damaged specimens were exposed to a specific timetemperature curve intended to produce a significant zone 

of damaged wood due to elevated temperatures within the 

residual wood section. The proposed test applies when flexural properties are to be refined beyond estimates that rely 

solely on reductions in cross-section and tabulated values 

for fire-damaged members and members that can otherwise 

be expected to remain in service.

A.2 Material Description

A.2.1 Design Values

The evaluations are based on tests of eight 24F-V8 DF/DF 

5-1/8-in. by 12-in. by 10-ft (130-mm by 305-mm by 3-m) 

glued-laminated beams. All beams were ordered from the 

same supplier and were used as the test specimens for this 

project.

A.2.2 Moisture Content and Conditioning

Upon arrival, each of the glued-laminated beams was conditioned until equilibrium moisture content to limit the 

variability amongst the beams because of moisture content 

(ASTM D 4933, ASTM 1999). The beams were conditioned 

using the heated-room dry method at a temperature of 83 °F 

(28.5 °C) with a 14.7% ± 1.0% relative humidity. Both the 

moisture content and relative humidity were tracked on a 

regular basis using a Delmhorst Moisture Check Meter and 

a Delmhorst Thermo-Hygrometer. All of the beams were 

then wrapped in Visqueen polyethylene sheeting to avoid 

reconditioning. Four of these beams served as the control 

group. These beams were not fire-damaged. The other four 

beams were fire-damaged in a furnace at FPL. 

A.2.3 Field Procedure and Equipment

The technique involved affixing a 6061-T651, type 200, 

aluminum block to the bottom of a beam with 1/4-20x2 

socket cap screws once 7/32-in. (5.6-mm) pilot holes were 

drilled into the bottom of the beam at these locations. Each 

screw was inserted to a torque of 30 lb-in. (3.4 N-m) at an 

insertion depth of 0.75 in. (19 mm). Next, increasing loads 

were applied and deflections were recorded. Both the beam 

and the aluminum block were subject to bending, although 

at different rates; as such, measurable changes in average 

torque readings resulted from the interaction between the 

aluminum block and bottom of the beam after the load was 

removed for the partial-composite load tests. 

The purpose of this was to establish a definitive target deflection that the beam could be reloaded to and that was 

within the linear elastic range. In so doing, the corresponding load, otherwise identified as the “40% estimate of 

maximum” load (hereafter referred to as the 40% load) is 

achieved once mean torque readings correspond to values in 

this range. Here, deflections were recorded along with values for the increasing, applied loads using a data acquisition 

system. The estimated maximum load is simply a multiplier 

of the 40% load. With this value, both the residual stiffness 

(apparent modulus of elasticity, Ef

) and the residual strength 

(modulus of rupture, SR) can be accurately estimated using 

conventional equations found in the appendix of D 198-09 

(ASTM 2009b). The difference is that the flexural property 

approximations now stem from a definitive load estimate in 

the linear elastic range. The tests performed on two expendable beams eliminated a considerable amount of trial and 

error that would have otherwise accompanied the load tests. 

The field equipment items include a digital level, a calibrated torque wrench, a pair of calipers, and a pre-drilled aluminum bar (attached to the bottom of a fire-damaged beam).

A.3 Laboratory Procedure

A.3.1 Fire Damage

Four beams were exposed to a specific time-temperature 

curve intended to produce a significant zone of damaged 

wood from elevated temperatures within the residual wood 

section. The 4-ft by 7-ft (1.2-m by 2.1-m) horizontal furnace 

has a series of eight diffusion-flame natural gas burners on 

the floor of the furnace. Each of the four 5-1/8-in. by 12-

in. by 10-ft (130-mm by 305-mm by 3-m) test beams was 

inserted in openings on the ends of the furnace and exposed 

to a relatively low fire exposure for a period of three hours. 

The tops of the beams were covered with gypsum board so 

three surfaces were exposed to the fire. Approximately 6 ft 

(1.8 m) of the middle of each beam was subjected to fire 

exposure. 

Exposure to the standard ASTM E 119 (ASTM 2010) fire 

exposure of fire-resistance tests would have resulted in a 

fairly steep temperature gradient within the wood beam. 

Since the intent was to evaluate the ability to determine 

loss in strength of the residual section, a longer and lower 

temperature fire exposure was used in hope of a more 

gradual temperature gradient within the charred member. 

Appendix—Development of New NDE Test for Fire-Damaged Wood
Evaluating Fire-Damaged Components of Historic Covered Bridges

31

A modification of a curve previously used to represent the 

plenum exposure in a protected wood truss assembly was 

employed (White and others 1993). This idealized time–

temperature curve is summarized in Table 1.

At the end of the approximately three hours of fire exposure, the beams were quickly removed from the furnace and 

sprayed with water. A series of thermocouples were inserted 

4 in. (102 mm) down from the top surface of the specimen 

at various distances of 0.25 to 2.5 in. (6 to 64 mm) from the 

surfaces of the sides of the beam. The temperature data were 

used to plot a temperature profile at the time of test termination. Analysis of the combined temperature data for all four 

tests indicated that the base of char layer with a temperature 

of 572 °F (300 °C) was approximately 0.3 in. (8 mm) inward. As was expected, the center temperature was much 

cooler at approximately 212 °F (100 °C). 

In contrast, the base of the char layer for the semi-infinite 

slab has a different temperature profile. Specifically, 356° F 

(180 °C) at 0.25 in. (6 mm) and 68° F (20 °C) at 1.3 in. 

(33 mm). When compared to a semi-infinite slab subjected 

to a standard fire-resistance test, the temperature was 491° F 

(255 °C) at 0.25 in. (6 mm) and was 252° F (122 °C) at 

1.3 in. (33 mm) inward.

There was more charring along the bottom of the glued-laminated beams than along the sides. Each glued-laminated, 

fire-damaged beam was cooled to room temperature prior to 

being packaged and shipped back to Montana Tech of The 

University of Montana for further evaluation and testing.

A.3.2 Laboratory Tests

After removing the char layer from each of the fire-damaged 

beams, the section properties for all beams were measured 

and recorded. The beams were then instrumented with epoxied (Vishay foil strain gauge) and removable Transducer 

Model BDI ST350 uniaxial strain gauges (Bridge Diagnostics, Inc., Boulder, Colorado). The two types of strain 

gauges were affixed at comparable locations, above, below, 

and near the depth-wise center of each beam.

Values collected from the gauges were used to verify the 

amount of composite action that was achieved. All beams 

were load-tested at Montana Tech of The University of 

Montana using a steel load frame (Fig. 12). The load frame 

consists of four HP Shapes, a reaction beam, and various 

support beams. Hydraulic loads were applied with a 10,000 

psi (69 MPa) capacity hydraulic hand pump and cylinder 

equipped with a tilt-saddle. The values associated with the 

applied loads and corresponding deflections were recorded 

using load cells and displacement transducers during the 

appropriate stages of testing (Fig. 12).

A.3.3 Creating Partial Composite Action between the 

Aluminum Stock and the Wood Beam

The results from the load-test data indicate that a modest, 

neutral axis shift of 2.5% or less was achieved. Further evaluation of these results indicated that the partial composite 

action between the aluminum stock and a glued-laminated 

beam was appropriate at this level. This is because the 

measurable changes in the mean machine screw torques 

were also germane to the proposed field test. The shift in 

the neutral axis of each beam was identified and the amount 

of partial-composite action was assessed from the strain 

measurements at three different locations. To measure this, 

values for strain were plotted at three distinct locations for 

three distinct linear-elastic loads. The intersection of the 

resulting lines that were fit to the data denotes the measured 

neutral axis. The neutral axes were measured with the aluminum stock affixed to the tension face of the beam (partialcomposite action) and without the aluminum stock affixed 

to the tension face (non-composite action). Fig. 12 depicts 

such a plot for a typical non-composite beam test.

The neutral axis shifts 2% on average when the partial-composite tests are compared to the non-composite tests using 

Table 1—Idealized time–temperature curve

Temperature Time

(°F) (°C) (min)

149 65 10

199 93 20

370 188 30

500 260 45

621 327 60

662 350 120

707 375 180

Figure 12—Neutral axis plot for a typical, non-composite 

beam.

–600 –400 –200 0 200 400 600

Non-composite

2

9

8

7

6

5

4

3
General Technical Report FPL–GTR–243

32

epoxied stain gauges. This shift is 2.5% on average when 

removable strain gauges are used. The results indicate that 

both the removable strain gauges and epoxied strain gauges 

are comparable. However, field applications may make one 

type of gauge preferable to the other.

A.3.4 Torque Readings

The purpose of the screw-withdrawal test is to establish a 

definitive, target deflection that individual beams can be 

re-loaded to that was within the linear elastic range. Torque 

readings of four different insertion depths were explored 

in this study: 0.75 in. (19 mm), 0.5 in. (13 mm), 0.25 in. 

(6.3 mm) and 0.125 in. (3.2 mm). The results from laboratory experiments and statistical analyses indicate that for 

both the control and fire-damaged group, the variability at 

an insertion depth of 0.25 in. (6.3 mm) was greater than the 

variability at insertion depths of 0.5 in. (13 mm) and 0.75 in. 

(19 mm). These results are based on the Brown-Forsythe 

test with an initial p-value <0.0001 when testing the equality 

of all variances for both groups. Note, at an insertion depth 

of 0.125 in. (3.2 mm), the variability of the machine screw 

torque values were unreliable because of missing values 

that occurred because the machine screws stripped out upon 

contact. 

The mean torque readings were compared using the 2-way 

ANOVA F-tests as well as the Tukey HSD multiple comparison procedure. Plots of the mean torque readings, along 

with individual 95% confidence intervals collected at these 

depths are depicted in Figure 13.

When comparing across the control and fire-damaged 

groups, as well as across insertion depths, the Tukey HSD 

multiple comparison procedure results, calculated at the 

95% experiment-wise error rate, are significantly different where the levels are not connected by the same letter 

(Fig. 13). Note that at the conclusion of the partial-composite load tests, residual machine screw depths of 0.25 in. 

(6.3 mm) or less are to be avoided because of an increased 

potential for variability in torque readings and (or) the inability to obtain them. We recommend that the average 

values for torque be used on a given beam and that the test 

is performed on a number of beams as a way of controlling for naturally occurring variability inherent to these 

measurements. 

We recommend that torque readings collected after a load is 

removed should conclude when torque readings fall between 

20 and 25 lb-in. (2.3 and 2.8 N-m) based on the results 

from the laboratory experiments. This recommendation is 

also supported through the results of the statistical analyses 

wherein the variability in torque readings is reduced as the 

insertion depth passes 0.25 in. (6.3 mm) and levels off at 

0.5 in. (13 mm). Torque readings within this range can be 

achieved by incrementally loading each of the beams in the 

laboratory and collecting torque readings once the load is removed. This range falls below the 95% confidence intervals 

for the average screw torque readings at insertion depths of 

0.75 in. (19 mm) and 0.5 in. (13 mm). As supported by the 

Tukey HSD procedure, any measurable changes in average 

torque readings at these levels is non-significant. Moreover, 

any measurable changes in average torque are attributed to 

the resulting interaction between the aluminum stock and 

bottom of the beam. It should also be noted that this range is 

above the 95% confidence intervals for mean screw torque 

readings at insertion depths of 0.25 in. (6.3 mm) and 0.125 

in. (3.2 mm). These values are also presented in Figure 13.

A.3.5 Calculation of the Modulus of Elasticity and the 

Modulus of Rupture 

The purpose of the screw-withdrawal test is to establish a 

definitive target deflection that individual beams can be 

re-loaded to and that was within the linear elastic range. In 

so doing, the corresponding load, otherwise identified as 

Figure 13—Summary of statistics for the control and firedamaged groups.

0.125

0

5

10

15

20

25

C

B A A

30

0.25 0.50 0.75

Fire damaged

0.125

0

5

10

15

20

25

C

B

A A

30

0.25 0.50 0.75

Control

Screw insertion 

depth (in.) Control Fire damaged

0.125 5.47, 8.09 4.22, 6.03

0.25 12.43, 15.94 12.05, 14.34

0.50 25.95, 27.14 26.48, 27.67

0.75 25.78, 26.97 26.59, 27.72
Evaluating Fire-Damaged Components of Historic Covered Bridges

33

the “40% estimate of maximum” load (hereafter referred 

to as the 40% load) is achieved once mean torque readings 

correspond to values in this range. (Here, deflections were 

recorded along with values for the increasing, applied loads 

using a data acquisition system.) The estimated maximum 

load is simply a multiplier of the 40% load. With this value, 

both the residual stiffness (apparent modulus of elasticity, 

Ef) and the residual strength (modulus of rupture, SR) can be 

accurately estimated, using conventional equations found in 

the appendix of D 198-09 (ASTM 2009b).

A.4 Results—Flexural Properties 

The estimated values for flexural properties were obtained 

by using the methodology established through the proposed 

field test. Table 2 presents the results for the torque tests, the 

corresponding target values for deflection, along with the 

actual and estimated flexural properties for all of the beams 

evaluated herein. 

The determination of the actual values for the beams tested 

destructively without the aluminum bar was made using the 

methodologies described in ASTM D 198 (ASTM 2009b) 

for each beam with solid rectangular homogeneous crosssection and center point loading. The modulus of rupture, 

SR, was calculated using the maximum bending moment at 

the maximum load borne by the beam. The apparent modulus of elasticity, Ef, was calculated using the equation in 

the appendix of D 198 with the load-deflection slope determined for the range of 20% to 40% of the maximum load. 

The data collected from the laboratory tests indicate that 

the averages for the estimated and actual Ef differed by less 

than 2% for the control group and less than 1% for the firedamaged group. Should beam L be omitted, the averages for 

the SR for the control group differ by approximately 10% 

where the actual values and estimated values are compared 

to one another. Including beam L, this average becomes 

larger. A comparison of averages made within the firedamaged group shows that the SR differs by less than 3% for 

the actual and estimated values. These results are depicted 

in Figure 14.

Table 2—Actual and estimated values for SR and Ef, all beams

Beam ID

Average 

torque 

reading,

lbf-in (N-m)

Target

deflection,

in. (mm)

Apparent Ef,

×103

 lbf/in2 (MPa)

SR,

lbf/in2

 (MPa)

Actual Estimate Actual Estimate

Control

Beam 3 24

(2.7)

0.327

(8.3)

749

(5,170)

793

(5,470)

5,060

(34.9)

4,640

(32.0)

Beam 7 24

(2.7)

0.339

(8.6)

736

(5,070)

729

(5,030)

6,740

(46.5)

5,530

(38.1)

Beam 8 20

(2.3)

0.218

(5.5)

1,109

(7,650)

1,106

(7,630)

5,000

(34.5)

4,900

(33.8)

Beam L 22

(2.5)

0.192

(4.9)

1,075

(7,410)

1,090

(7,520)

7,440

(51.3)

4,150

(28.6)

Average — — 917

(6,320)

930

(6,410)

6,060

(41.8)

4,810

(33.1)

Fire damaged

Beam 2 24

(2.7)

0.247

(6.3)

810

(5,590)

818

(5,640)

3,620

(25.0)

3,600

(24.8)

Beam 4 22

(2.5)

0.306

(7.8)

645

(4,450)

652

(4,500)

3,890

(26.8)

3,750

(25.9)

Beam 5 25

(2.8)

0.145

(3.7)

887

(6,120)

891

(6,150)

2,560

(17.6)

2,440

(16.8)

Beam 6 23

(2.6)

0.277

(7.0)

570

(3,930)

567

(3,910)

3,060

(21.1)

2,990

(20.6)

Average — — 728

(5,020)

732

(5,050)

3,280

(22.6)

3,190

(22.0)
General Technical Report FPL–GTR–243

34

A.5 Conclusions

This research presents a proposed field technique that can 

be used to determine the residual flexural properties in 

glued-laminated beams that have been fire-damaged. Two 

treatment groups were established: a control group and a 

fire-damaged group. The nondestructive values, arrived 

at through the proposed field test, were compared to the 

destructive values for the beams from both groups. The following conclusions are based on these results.

Results show that on average the relative changes in the 

apparent modulus of elasticity and the modulus of rupture 

changed by 20% and 46% after the beams had been firedamaged. This marked decrease in flexural properties goes 

beyond simply reducing the cross-section once the char 

layer has been removed. 

Results of this research indicate that torque readings collected after a load is removed should conclude when torque 

readings fall between a range that is inclusive of 20 and 25 

lb-in. (2.3 and 2.8 N-m). The same finite range was applicable to beams in either group. This approach has been validated in the laboratory and is specific to the beams tested 

herein.

Results corroborate the proposed field technique when direct 

determinations of estimates of the residual flexural properties are sought. This is important where like members that 

exhibit various degrees of degradation are tested in the field 

and are otherwise expected to remain in service because the 

flexural property approximations now stem from a definitive 

load estimate in the linear elastic range.

2

4

6

8

10

12

0

2 4 6

Estimated values

SR × 1,000 lb/in2

Ef × 10,000 lb/in2

Equality

Actual values

0 8 10 12

Figure 14—Estimated and actual values for SR and Ef, all beams.