The NIST 3 Megawatt Quantitative 

Heat Release Rate Facility 

Rodney A. Bryant

Thomas J. Ohlemiller

Erik L. Johnsson

Anthony Hamins

Brian S. Grove

William F. Guthrie

Alexander Maranghides

George W. Mulholland

NIST Special Publication 1007

0 1000 2000

0

1

2

Heat Release Rate, MW

Time, s
NIST Special Publication 1007

The NIST 3 Megawatt Quantitative 

Heat Release Rate Facility 

Rodney A. Bryant

Thomas J. Ohlemiller

Erik L. Johnsson

Anthony Hamins 

Building and Fire Research Laboratory

Brian S. Grove

Department of Justice

Alcohol Tobacco and Firearms

Fire Research Laboratory

William F. Guthrie

Information Technology Laboratory 

Alexander Maranghides

George W. Mulholland

Building and Fire Research Laboratory

December 2003

U.S. Department of Commerce

Donald L. Evans, Secretary

Technology Administration

Phillip J. Bond, Under Secretary for Technology

National Institute of Standards and Technology

Arden L. Bement, Jr., Director
Certain commercial entities, equipment, or materials may be identified in this 

document in order to describe an experimental procedure or concept adequately. Such 

identification is not intended to imply recommendation or endorsement by the 

National Institute of Standards and Technology, nor is it intended to imply that the 

entities, materials, or equipment are necessarily the best available for the purpose.

National Institute of Standards and Technology Special Publication 1007

Natl. Inst. Stand. Technol. Spec. Publ. 1007, 81 pages (December 2003)

CODEN: NSPUE2

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 2003

_________________________________________

For sale by the Superintendent of Documents, U.S. Government Printing Office

Internet: bookstore.gpo.gov — Phone: (202) 512-1800 — Fax: (202) 512-2250

Mail: Stop SSOP, Washington, DC 20402-0001
iii

EXECUTIVE SUMMARY 

The 3 Megawatt Heat Release Rate Facility (3MWHRRF) was developed at NIST as a 

first step toward having broad capabilities for making quantitative large scale fire 

measurements. Such capabilities will be used at NIST to validate fire models and to 

develop sub-grid models. It will also serve to provide a data base for studying a broader 

range of fire phenomena, and to address issues related to material acceptance and fire 

codes. An equally important objective is to provide templates for use by other 

laboratories including commercial testing facilities to improve the quality of their data. 

Heat release is the result of the combustion of a fuel with the oxygen in air. The fuels of 

primary interest are those found in constructed facilities and include wood, plastics, foam 

materials used in furnishings (such as polyurethane), wire insulation (such as polyvinyl 

chloride), and carpet materials (such as nylon). 

The rate at which heat is released is the single most important quantity in terms of fire 

safety. Thus it is important that this measurement be made in a quantitative manner. It is 

a key predictor of the hazard of a fire; directly related to the rate at which heat and toxic 

gases build up in a compartment or the rate at which they are driven into more remote 

spaces. Heat release rates on the order of 1 MW to 3 MW are typical in a room that is 

flashed over or from a single large object such as a bed or sofa. 

It is important that heat release rate measurements be made accurately because fire 

regulations are frequently based on peak rates of heat release. Testing laboratories must 

be confident that the objects tested pass the required regulation and manufacturers need 

accurate information in defining the fire safety characteristics of their products. A second 

need for accurate heat release rate data is for the development of quantitative models for 

predicting heat release rate. In comparing a fire experiment and a model prediction, it is 

essential that the heat release rate measurement have an estimated uncertainty. 

The 3MWHRRF developed at NIST meets the needs described above for objects that can 

be placed under the 6 m × 6 m hood, which is approximately 4 m above the floor, or for 

enclosures whose effluent can all be directed into the hood. It is capable of measuring 

heat release rates in the range of 0.10 to 3.0 MW including brief peaks as high as 5 MW. 

As documented in this publication, the expanded uncertainty (95 % confidence interval) 

is 11 % of the heat release rate for fire sizes larger than 400 kW. The response time of 

the system is such that it can accurately resolve dynamic heat release rate events of 15 

seconds or more. 

This document is intended to serve as a description of the NIST 3 Megawatt Heat Release 

Rate Facility and as a general guide for implementing, operating and maintaining quality 

control of similar quantitative large scale heat release rate measurement facilities. The 

measurement system consists of a hood and duct system, a duct flow rate measurement, a 

gas flow sampling system, and a gas analysis system. Key features of the design and 

fabrication of these systems are presented including detailed schematics and the rationale 
iv

for the choices made. Transient response and gas dispersion in the sampling system, 

sampling from a non-uniform duct flow, and making accurate total flow measurements 

using bi-directional probes are dealt with in the design. 

A calibration burner is essential for the operation of the 3MWHRRF and for 

characterizing its time response. The design of the natural gas burner including the 

special flow control and flow metering capabilities are presented. 

The general design of the software allowing the display of nearly real-time heat release 

rates with corrections for sampling/instrumental delay and response times, is described. 

The uncertainty analysis described in the report includes the effect of the drift in the gas 

analyzer outputs, the correlation effects in computing the oxygen difference relative to 

the background value, and combining the uncertainties from over 50 variables in the heat 

release equation. 
v

ACKNOWLEDGEMENTS 

The success of any project of this scope and magnitude is due to more than just the work 

of the authors. Therefore the authors wish to thank the following NIST staff for their 

valuable contributions: Laurean A. DeLauter, David W. Stroup, Roy A. McLane and 

Gary Roadarmel for construction of the calorimeter, and William H. Twilley, Laurean A. 

DeLauter, and Jay A. McElroy for construction of the natural gas calibration burner. 
vi

TABLE OF CONTENTS 

page

EXECUTIVE SUMMARY III

ACKNOWLEDGEMENTS V

1 BACKGROUND 1

2 DESCRIPTION OF SYSTEM 5

2.1 CALORIMETER 7

2.2 DATA ACQUISITION SYSTEM 15

2.3 CALIBRATION BURNER SYSTEM 19

2.4 DESIGN CONSIDERATIONS 22

3 OPERATION AND SAFETY 40

3.1 CALORIMETER 40

3.2 DATA ACQUISITION SYSTEM 40

3.3 CALIBRATION BURNER 41

3.4 SAFETY 42

3.5 BASELINE HEAT RELEASE RATE 42

4 ASSESSMENT OF UNCERTAINTY 45

4.1 TYPE A UNCERTAINTY 46

4.2 TYPE B UNCERTAINTY 47

4.3 UNCERTAINTY PROPAGATION 49

4.4 REPEATABILITY, EXPANDED UNCERTAINTY, AND CONFIRMATION COMPARISON 60

5 INTERNAL QUALITY CONTROL PROCEDURES 63

5.1 CERTIFIED OPERATOR 63

5.2 CALIBRATION OF INSTRUMENTS 64

5.3 CALIBRATION RECORDS 65

5.4 TRAPS AND FILTERS 65

5.5 DATA ACQUISITION 66

5.6 DATA ARCHIVING 66

5.7 CONFIRMATION OF HRR FACILITY 66

6 CONCLUDING REMARKS 67

REFERENCES 69

APPENDIX A NOMENCLATURE 72

APPENDIX B UNCERTAINTY COMPARISON EXAMPLE 73

B.1 UNCERTAINTY IN THE DIFFERENCE OF CORRELATED MEASUREMENT QUANTITIES 73
1

1 Background 

Heat release rate is defined as the enthalpy change per unit time as a result of the 

conversion of the chemical energy of a fuel to heat in a combustion process. Most 

commonly the fuel is carbon-based and the combustion process is one of oxidation, 

usually by the oxygen in air. Heat release rate is typically reported in kilowatts (kW) or 

megawatts (MW). Its importance follows from the fact that it is a key predictor of the 

hazard of a fire, directly controlling the rate at which heat and toxic gases build up in a 

compartment or are buoyantly driven into more remote spaces [1]. It is, in effect, the 

most pertinent measure of the size of a fire [2]. 

Understanding the heat release rate potential of an object or an assembly of objects is 

essential to estimating the hazard such an object or arrangement could pose if ignited. 

Rational limitation and regulation of such a hazard is based on such understanding. 

Although substantial progress is being made, the present state of fire growth models is 

such that accurate prediction of heat release rate for common room contents such as 

chairs or beds cannot be made with full confidence. Thus measurements of heat release 

rate on real objects are currently essential. The measured results may be used as input for 

hazard calculation models such as CFAST (The Consolidated Model of Fire Growth and 

Smoke Transport) [3] or for product regulation. In either case, accurate heat release rate 

measurements are essential to meaningful results. 

For a simple object made from a pure substance with a known constant heat of 

combustion the heat release rate measurement requires only a mass loss rate measurement 

on the object since the former is proportional to the latter, i.e. 

q = m∆Hc & & ( 1 ) 

where q& is the measured rate of heat release, m& is the mass loss rate (time derivative of 

decreasing object mass) and ∆Hc is the known heat of combustion (e.g., kJ/kg). In 

practice the heat of combustion of a complex object is rarely tabulated and in any case is 

variable for charring materials. Thus reliance on mass loss to quantify heat release is 

unreliable for real objects. On the other hand, mass loss is useful if the “object” is a welldefined gas or a liquid supplied at a known rate. This is, in fact, the basis for calibration 

of more general heat release rate calorimeters, where the calorimeter output can be 

compared with a precisely measured flow of a fuel whose heat of combustion is also 

accurately known. 

Early measurements of heat release rate focused on the heat output as represented by the 

temperature of the product gases. Thus an array of thermocouples was positioned in a 

duct which captured all these gases to measure their average gas temperature [ ASTM 

1321 (LIFT test); ASTM 906 (OSU calorimeter)]. This approach assumes the flow is 

adiabatic, which is never a reality. The heat emitted from the fire as radiation, for 

example, can be 35 % or more of the total chemical heat release rate. More sophisticated 
2

variants of this thermal approach, based on measuring the fuel gas flow rate required to 

maintain a constant product stream temperature, had the same limitation [4]. 

Huggett [5] first suggested a very different approach, now termed oxygen consumption 

calorimetry.1

 He followed up on the much earlier finding that the amount of heat evolved 

from most organic materials per unit mass of oxygen consumed in their complete 

combustion is nearly constant. For example, for butyl alcohol: 

_ 2

4 9 2 2 2 2 2

( ) 12.79 /

6 6(3.76) 4 5 6(3.76)

2 H MJ kgO

C H OH O N CO H O N

∆ c Mass O = −

+ + → + +

( 2 ) 

Huggett shows that for a wide variety of molecules from pure hydrocarbons, to partially 

oxygenated species, to partially halogenated species, and a wide variety of polymers and 

natural materials such as wood and coal, the heat release per unit mass of oxygen falls 

within a narrow range. He showed that, for most common materials containing C, H, O, 

and N, the average heat release per unit mass of oxygen is 13.1 MJ/kg O2; the standard 

deviation of the heat release rate for materials examined by Hugget is 0.35 MJ/kg O2. 

Thus the oxygen deficit in the duct flow (relative to ambient air) is a measure of the heat 

release rate in the flow which the duct captures. 

Sensenig built the first working apparatus based on oxygen consumption calorimetry [7]. 

The heat release rate was inferred from the measured oxygen deficit in the flow of a hood 

that captured the fire plume as shown in the following equation: 

θ o Hc Vol O VD X O q _ 2 2 ( ) & & = ∆ ( 3 ) 

where o

O O

o X O X X 2 2 2 θ = ( − )/ ( 4 ) 

with o X O2 being the oxygen volume fraction in the ambient air, O2 X the oxygen volume 

fraction in the duct after the captured smoke plume (and any excess air) is well-mixed, 

VD

& is the measured volume flow rate in the hood duct, and _ 2 ( ) ∆Hc Vol O is the heat of 

combustion per unit volume of oxygen consumed (e.g., MJ/m3

 O2). This simplified 

equation is approximate since it neglects other gases that are present such as water vapor 

and CO2. The equation does convey, however, the role of the primary variables: duct 

flow rate and duct oxygen depletion. 

Equation (3) may be rewritten on a mass basis as follows: 

( ) ( ) 2 2 2 _ O

o q& = ∆Hc Mass O m& O − m& ( 5 ) 

 

1

 Factory Mutual also uses a related approach in which the production of carbon dioxide and carbon 

monoxide, rather than the loss of oxygen, are taken as a measure of heat release [6] 
3

where _ 2 ( ) ∆Hc Mass O is the heat released per unit mass of oxygen, o mO2

& is the mass flow 

rate of oxygen in the duct prior to the fire test, and O2 m& is the mass flow rate of oxygen in 

the duct during the fire test. 

For an open system only the exhaust flow rate, which includes the flow rate of incoming 

air and combustion products, is measured. It is necessary to relate the measured exhaust 

flow rate to the incoming air flow rate. Therefore it is useful to define the oxygen 

depletion factor, φ , as the fraction of incoming air that is depleted of its oxygen. 

o

O

O

o

O

m

m m

2

2 2

&

& − & φ = ( 6 ) 

In relating the mass flow of air into the system to the mass flow in the duct, the 

assumption is employed that nitrogen is not involved in the combustion process and 

therefore the nitrogen flow is constant. Nitrogen analysis is not performed so its volume 

fraction is replaced using the assumption that only N2, O2, CO2, CO and H2O are present. 

This substitution is most accurate if all of the other gas species are measured. Accurate 

water analysis is difficult, so it is typically approximated on the basis of the humidity of 

the incoming air. Using the substitution for nitrogen, φ is expressed in terms of the 

measured volume fractions of the gas species. 

o

O CO CO O

o

CO CO O CO

o

O

X X X X

X X X X X

2 2 2

2 2 2 2

(1 )

(1 ) (1 )

− − −

− − − − φ = ( 7 ) 

It is apparent in Eq. (2) that the combustion process of hydrocarbon fuels can result in a 

greater number of moles of products than reactants. Parker [8] defines a chemical 

expansion factor as follows: 

1 ( 1) 2 α = + β − o X O ( 8 ) 

where β is the ratio of the moles of combustion products to the moles of oxygen 

consumed. In the case of hydrocarbon fuels the average value of α is 1.10. 

The number of moles in the fraction of air depleted of its oxygen is replaced by an equal 

or greater number of moles of combustion products (second term of the equation below). 

Therefore, the mass flow rate of air into the system, mair & , may be related to the measured 

mass flow rate in the exhaust, me & , by the following relation: 

φ φα

air

air

air

air

e

e

M

m

M

m

M

m& & & = (1− ) + ( 9 ) 
4

Assuming that the molecular weight of the exhaust duct gas is approximately equal to the 

molecular weight of the incoming air (Me ≈ Mair), the mass flow rate of the incoming 

oxygen is expressed as: 

air

o O

O inc

e

air

o O

air O inc

o

O M

M

X m

M

M

m m X 2

2

2

2 2 , , 1+ ( −1) = = φ α

& & & ( 10 ) 

Paramagnetic oxygen analyzers require the water to be removed from the sampled gas, 

therefore a correction must be made to the ambient oxygen volume fraction 

measurement. 

o

O

o

H O

o X O inc X X 2 2 2 (1 ) , = − ( 11 ) 

Using Eq. (6), (10), and (11), Eq. (5) may be rewritten as follows with φ , as described in 

Eq. (7), representing the case where O2, CO2, and CO are measured while H2O is 

removed from the sampled gas: 

air

o O

O

o

H O

e

c Mass O M

M

X X mq H 2

2 2 2 (1 ) 1 ( 1) ( ) _ − + − = ∆

φ α

φ & & ( 12 ) 

Incomplete combustion, especially CO formation, often occurs in fires and CO formation 

yields only about half as much heat per unit mass of O2 consumed as does CO2

formation. Huggett points out that even if the CO concentration is 10 % of the CO2

concentration, the effect on the net heat released per mass O2 consumed is small. 

However, flame retardants frequently boost CO formation and oxygen-starved burning, 

as in flashover conditions, does so as well. Often CO formation may exceed 10 % of 

CO2 formation. In general, it is recommended to measure CO and correct the heat release 

rate for its formation as shown in the following equation: 

air

o O

O

o

H O

e

O

HC CO

c Mass O

CO

c Mass O

HC

c Mass O

M

M

X X m

X

X

q H H H

2

2 2

2

2 2 2

(1 ) 1 ( 1)

] 2

1 [( ) (( ) ( ) ) _ _ _

− + −

− = ∆ − ∆ − ∆

φ α

φ φ

&

&

( 13 ) 

where 

HC Hc Mass O2 _ (∆ ) = heat of combustion of hydrocarbon fuel, MJ/kg O2

CO Hc Mass O2 _ (∆ ) = heat of combustion of carbon monoxide, MJ/kg O2

φ = oxygen depletion factor 

me & = mass flow rate in exhaust duct, kg/s 

α = combustion products expansion factor 

Mi = molecular weight of gas i, kg/kmole 

Xi = volume fraction of exhaust gas i 
5

o Xi = volume fraction ambient gas i 

The algebraic complexity of Eq. (13), compared to Eq. (3) or (5), arises essentially from a 

more thorough treatment of the mass balances (and atom balances) on the species 

between the fire, where they are generated, and the instruments, where they are 

measured. Parker [8] and Parker and Janssens [9] discuss the details of the heat release 

rate calculation based on the extent to which the duct gas flow is characterized, i.e., 

whether one measures only oxygen or also includes CO, CO2 and water. The NIST 

calorimeter system measures oxygen, CO and CO2; a correction is made for ambient 

humidity as an estimate of the water concentration in incoming air. The heat release rate 

is not corrected for the soot yield2

. Details concerning the input measurement quantities 

for Eq. (13) are discussed in Sec. 4, Assessment of Uncertainty. 

It is important to realize that the oxygen deficit in the duct flow, even in large fires, is 

typically rather small. This is a consequence of the need to prevent overheating of the 

exhaust duct and to allow completion of the combustion reactions before the fire plume 

enters the duct. Thus the best one can do with regard to capturing the plume is draw it 

and no other diluting air into the duct inlet at the top of the hood [10]. However, fire 

plumes are typically already quite diluted at the flame tips [11] and, as a result, one can 

expect an oxygen reduction only from 21 % by volume to about 19 %. Rarely is it 

actually possible to capture the plume in this manner, so additional plume dilution is the 

norm. An implication of this is that the oxygen meter needs to have high resolution if the 

inferred heat release rate is to be precise. Suffice it to say that oxygen measurement and 

duct flow rate are the two measurements in need of greatest attention. 

2 Description of System 

The 3 MW Heat Release Rate Facility is located in the Large Fire Research Facility and 

became operational in August, 2001. It employs oxygen consumption calorimetry in the 

standard manner to measure the total heat output of a fire. The description of the facility 

may be segmented according to three major functions: calorimetry, data acquisition, and 

operational validation. Each of these functions is shown as boxes in Figure 1. Within 

each box are the subsystems that combine to complete the function. To summarize the 

functions, calorimetry consists of collecting all of the gases produced by the fire, 

measuring the rate of flow of the gases, sampling a small and well mixed portion of the 

gases, and measuring the volume fractions of specific gas species (especially oxygen 

volume fraction). Data acquisition consists of sampling the electrical output of each 

calorimetry instrument as it responds to the physical conditions that it monitors, 

converting the electrical output to a digital signal, recording a time history of the digital 

signal and converting the record to a value of scientific significance and displaying it in 

near real-time. The data acquisition system also provides a user interface to control the 

 

2

 A first order thermodynamic analysis indicates that for highly sooting toluene flames the effect on the 

heat of combustion is less than the uncertainty in the heat of combustion per kg O2 for an unknown fuel. 
6

output of the calibration burner and record the subsequent response of its instruments. 

Operational validation consists of putting a controlled flow of chemically released heat 

into the system and measuring the heat flow independent of the calorimeter measurement. 

The calorimeter and data acquisition systems are essential for measuring the heat released 

from a burning item. The calibration burner is essential for confirming the measurement. 

Detailed descriptions of each system and selected subsystems are presented in the 

sections that follow. 

Exhaust System

Calibration

Burner

Duct Flow

Measurement

Gas Analysis

System

Volume Flow

Control

Volume Flow

Measurement

Specified 

Heat Input

Unknown

Heat Input

or

Calorimetry

Validation

Gas Sampling

System

Signal 

Conditioning,

A/D and D/A

Conversion

Data Acquisition

Figure 1 Schematic overview of the 3 MW Heat Release Rate Facility. 

The system is nominally rated at 3 MW total heat release rate but fires with brief peaks as 

high as 5 MW have been successfully measured. (The hood skirt provides a temporary 

buffer capacity, as discussed below.) Calorimeter capacity may be limited either by the 

ability of the hood to capture fire plumes (i.e., by the total volumetric flow capacity of the 

hood) or by the height of the flame tips. The user must have a reasonable idea 

beforehand whether either of these factors will make this hood suitable for a specific fire 

situation. The literature and/or experience is the only guide as to probable maximum fire 

size. Flame tip height can be calculated from the available pool fire correlations [11] 

(and some results of this type are included below) but the user should be aware that not 

all fires behave like simple pool fires. The vertical extent of a burning object can affect 

flame tip height and plume entrainment, as can a tendency for wide objects to burn with 

more than one fire plume. (Multiple plumes may entrain more air than a single plume 

involving the same total heat release rate.) Reference [11] also includes formulas for the 

calculation of plume entrainment and thus the total volumetric or mass flow which a hood 

must swallow; again these are based on pool fires. 
7

Fires as small as 50 kW can be measured in the 3 MW calorimeter, if a degraded signal to 

noise behavior is acceptable. Objects giving peak heat release rates this small are best 

measured in a smaller calorimeter hood. 

2.1 Calorimeter 

2.1.1 Hood and Duct System 

Side and top view diagrams of the 9 m x 12 m and 6 m x 6 m hoods and duct systems are 

shown in Figure 2. The test area beneath the 6 m x 6 m hood is open on three sides. The 

fourth side is relatively close to one wall of the building; this could impact air inflow 

symmetry, especially for fires larger than 2 MW, though no problems have been noted to 

date. The base of the hood is located 4.57 m above the laboratory floor. The effective 

height of the hood inlet has been lowered to 3.0 m by the addition of a vermiculite glass 

fiber skirt (rated for a maximum temperature of 923 K)3

. The skirt (in addition to the 

1.80 m depth of the hood itself) offers a buffer volume for plume accumulation in the 

event of a transient peak in fire behavior that yields a plume flow too great for the hood 

to swallow. It can also serve to limit plume entrainment for fires that persist in exceeding 

the rated capacity of the hood. Both roles come at a price, as discussed below in the 

Design Considerations section.4

 The effectively lowered hood inlet also helps assure that 

a weakly buoyant plume from a small fire is fully captured before it can drift out from 

under the hood. 

The interior surface of the hood is coated with a fibrous cement mix to provide thermal 

insulation of the steel structure. This is an essential function, particularly to 

accommodate large fires that persist tens of minutes. However, for large and sustained 

fires, heating of the insulation material (and of the hood skirt) could cause it to become a 

source of re-radiation back to the burning object, accelerating the rate of burning. 

A 1.52 m inside diameter steel duct emerges from the top of the hood and immediately 

turns to become horizontal, running under the high bay roof (Figure 2). Immediately 

downstream of this are inlets for ducts from smaller calorimeter hoods. These are 

normally sealed off to preclude dilution of the captured plume gases from the six-meter 

hood.5

 The duct then makes a right angle turn upward, passes through the roof and 

makes a second right angle turn to run along the top of the roof. The entire length of the 

duct is thermally insulated to minimize condensation of plume moisture. 

 

3

 Above the fiber cloth is a lighter layer of fiberglass material used to seal the gap between the fiber cloth 

(hanging from a cable) and the interior surface of the hood. This seal is necessary to prevent significant 

escape of smoke in some circumstances. 

4

 In general it is much better to use a skirt than to risk losing part of the fire plume. When part of the plume 

is lost, the oxygen deficit generated by the fire is underestimated and thus so also is the heat release rate. 

Quantifying this is very difficult since one cannot judge the fraction of a plume that is being lost unless it is 

very obviously negligible. 

5

 Small inward leaks of air at this point in the duct system should not be problematical since they would 

have time to mix over the duct cross section before the gases are sampled for calorimetry. Such leaks are 

analogous to fire plume dilution, however, and the larger they are the more they lower the oxygen deficit in 

the duct and thus the system sensitivity to heat release rate. 
8

0.91 m

Gas Sampling Plane

Velocity Measurement Plane

Flow

Straightener

D1.98 m

D1.98 m DN Thru Roof Gas Sampling Plane

Velocity Measurement Plane

D1.52 m

Flow

Straightener

D1.52 m DN Thru Roof

Top View

Above Roof

Access Hatch Access Hatch Access Hatch

9.00 m x 12.00 m Hood

6.00 m x 6.00 m Hood

4.57 m

6.40 m

1.83 m

D1.98 m

MOD

MOD

11.43 m

1.32 m D1.52 m

Figure 2 Large Fire Research Facility exhaust hoods and duct system. 

The last turn includes a dead end duct section (with an access hatch) that has been sealed 

off with a flat bulkhead having small holes to assure pressure equilibration. Dead 

volumes in the flow system such as this, not cleanly swept by flow, are sources of 

dispersion or spread of the time history of the heat release rate information carried in the 

plume flow and should be avoided throughout the duct system and the entire gas 

sampling system. They provide pockets of gas that are re-distributed into the duct flow at 

times later than the time at which they originally reached the pocket location. This tends 

to smear out heat release rate peaks and slow the system response time. There are 

dispersive processes of this nature occurring at various locations in any real flow and gas 

analysis system and minimization of these must be factored in to the system design. A 

more extensive discussion of those pertinent to this system is given below in Sec. 2.4.3. 
9

The three turns plus turbulence in the duct serve to mix the flow reasonably well. The 

mixedness of the flow was checked with helium injected at the hood duct inlet, sampled 

downstream and measured with a helium detector. These results indicated a uniform 

concentration over the sampled cross-section to within ± 5 %. The sampled section of the 

duct was the same section as that used for normal heat release rate measurements. 

The turns in the duct leave the time-averaged velocity profile skewed toward the top of 

the duct6

. There is not sufficient duct length available on the roof to guarantee a fullydeveloped turbulent flow profile after the last turn in the duct. A flow straightening 

section was added between the last turn and the gas velocity measurement section to 

reduce the eddy size and remove any swirl. The flow straightener is of a simple tube 

bundle design [12]; it consists primarily of 10 cm diameter thin wall steel pipe sections, 

1.5 m long. These are stacked in the section indicated in Figure 2 so as to fill the entire 

duct cross section. While the pressure head loss it induces is insufficient to fully flatten 

the velocity profile, measurements of the velocity at six locations across the duct cross 

section are sufficient to determine the total flow, as described in the next section. 

Approximately 2.3 m downstream of the gas flow characterization section, the duct 

expands asymmetrically to an approximate 1.98 m inside diameter, just prior to joining 

with the duct from the 9 m by 12 m hood. To minimize any effect of this asymmetry, the 

gas velocity measurement plane in the 1.52 m diameter section was kept 1 1⁄2 diameters 

upstream of the duct expansion section. 

Water condensation in the duct is a possibility in cold weather when humid lab air, or 

even more so, smoke, contacts the duct surfaces in the area atop the roof. If the duct 

walls are below freezing (there is no heating system in the duct), condensed water will 

freeze. Ice on the duct walls is probably of little concern. Ultimately the warm air will 

heat up the duct walls to the point that any ice melts and the water will then re-evaporate. 

Thermocouples on the duct wall allow monitoring of its temperature. 

2.1.2 Duct Flow Rate Measurement 

The velocity is measured approximately 4.2 duct diameters downstream of the outlet of 

the first flow straightener (located in the 1.52 m diameter section of the duct). Six bidirectional pressure probes of conventional design [13] are used. They are arrayed 

vertically on a duct diameter7

 (in one plane) and placed at the Tchebyachev locations 

recommended in Ref. [14] (see Figure 3). The six velocity measurements are equally 

weighted in calculating the total flow rate in the duct. The temperature near each velocity 

probe location is measured with a 1.6 mm diameter sheathed chromel/alumel 

thermocouple (one thermocouple per velocity probe). The velocity probes are held in 

place by a rigid, 13 mm wide, vertical steel support spaced about 15 cm away from the 

 

6

This skewing was predicted by LES modeling of the flow and was confirmed by actual mapping with a 

probe inserted from the top of the duct at the normal velocity measurement cross section. 7

 The implicit assumption of bi-lateral flow symmetry appears justified by the lack of left/right asymmetry 

in the system before this plane and by the probing of the flow in this region. 
10

probes to minimize any impact its cross section may have on the local velocity around the 

probes. The thermocouples are wound around the lead tubes of the bi-directional probes. 

Gas Sampling

Tee (holes face

downstream)

Bi-Directional 

Probes

Probe Leads

Probe Support

4.83 cm

20.4 cm

48.4 cm

4.83 cm

20.4 cm

48.4 cm

30.0 cm

Front View Side View

Probe Support

Flow

Gas 

Sampling

Tee

Tube to 

Heated Hose

(not shown

in front 

view)

Figure 3 Schematic of velocity probes and gas sampling tee mounted in the exhaust 

duct. The size of the gas sampling tee is exaggerated in order to clarify the location 

of the sampling holes. 

External to the duct and shown in Figure 4, the bi-directional probe lead tubes are 

bundled together and insulated with polyurethane foam pipe insulation (covered with 

aluminum tape for weathering) to minimize any temperature differential in the lead pairs 

from any given probe, since this can cause a zero offset and/or zero drift via differing gas 

column densities in the vertical portions of the probe leads8

. 

The bi-directional probe leads and their associated thermocouple leads go directly into an 

air-conditioned enclosure (see Figure 4) sitting on the building roof, next to the gas 

characterization section of the duct. Temperature control is necessary to minimize drift 

in the output of the six pressure transducers (MKS Baratron 220; range 0 to 133 Pa) that 

read the pressure differentials across the six bi-directional probes. 

The six pressure transducers used here have all been calibrated using a Dwyer Model 

1430 point gage manometer. This is a simple device that allows accurate imposition of 

the very low pressures seen in the duct flow velocity range that is pertinent here (less than 

100 pascals). The calibration results confirmed the factory-supplied data. 

 

8

 This is potentially a significant source of zero drift in the pressure signal as the sun moves across the sky; 

thus the thermal shielding is necessary. 
11

At a volumetric flow of 10.8 m3

/s, the velocities range from about 4 m/s to about 8 m/s. 

The local velocity values have fluctuations of roughly ± (2 to 5) % for this case, 

presumably due to turbulence in the duct. 

Enclosure Probe Leads

Figure 4 Exhaust duct velocity measurement and gas sampling station. 

2.1.3 Gas Flow Sampling System 

The gas in the duct is sampled approximately 30 cm downstream of the plane of the bidirectional probes. The sampler is a four-armed cross that spans the duct diameter 

vertically and horizontally, as shown in Figure 3. The ends of the arms are anchored to 

the duct wall. The arms are 25.4 mm OD tubing, each having four 6 mm diameter holes 

spaced at equal duct area points along the arm span. The holes face downstream to lessen 

the chances of their being plugged by soot or ash drawn onto the hood, and they are 

located on equal annular area radii. The gas is pulled through the center of the cross and 

then goes into a lead tube that passes out the bottom of the duct. The line from the exit of 

the duct to the pump enclosure is a 12.7 mm ID Teflon-lined heated hose, typically 

operated at 363 K. 

Suction is provided by an Air Dimensions Inc. Model 19320 TM heated double-head 

diaphragm pump that delivers 36 L/min at 35 kPa above atmospheric pressure when run 

at room temperature. When run at 373 K, it is expected to deliver about 20 % less 

volumetric flow (at NTP conditions), since it is a constant volume flow device. This 
12

pump is mounted in a separate enclosure beneath the enclosure containing the pressure 

transducers. The pump is protected from soot deposition with a filter on the inlet side. 

The flow rate of the gas sample is much more than is needed by the analytical 

instruments; together they require less than 5 L/min at one atmosphere. The reason for 

the excess gas flow is to shorten the time for getting the sample to the control room. The 

time to go from the gas sample tee to the control room is roughly 3 s. 

Condensation of water (especially) and any other volatiles in the sampled gas is a 

concern, especially in cold weather. The heads on the gas sampling pump are heated to 

minimize condensation. The pump enclosure is also heated. The line leading from the 

pump enclosure to the control room is 9.5 mm ID by 19.5 m long, Teflon-lined, heated 

tube. The line is typically run at a set point temperature of 348 K. 

Both the pump enclosure and the pressure transducer enclosure are mounted with wire 

rope vibration isolators (Enidine, Inc., Orchard Park, NY) between the enclosures and the 

adjacent supports. The principal reason for this is the sensitivity of the pressure 

transducers to any form of vibration. Both the gas sampling pump and the burners in the 

pollution control system can be significant sources of vibration of relatively low 

frequency (below about 30 hertz). The wire rope isolators offer better low frequency 

isolation than do rubber mounts. 

2.1.4 Gas Analysis System 

Figure 5 is a photograph of the flow control panel for the gas analysis system and Figure 

6 shows a complete diagram of the gas analysis system located in the control room. As 

soon as the gas arrives at the analysis system, much of the excess flow is dumped to an 

exhaust line.9

 The excess flow that is being dumped at the system inlet can be read with 

the large flowmeter. The flow normally bypasses this meter to minimize deposition of 

soot that may have passed through the pump filter. When the valves are switched so as to 

send the dump flow temporarily through the flowmeter one can check that the roof pump 

is behaving normally. The meter should read full scale (42.5 SLM) when the roof pump 

and the heated transfer lines are at their normal operating temperatures. 

The pressure in the region where the flow from the roof enters the gas analysis system is 

normally kept at 500 Pa above atmospheric pressure to assure positive flow out through 

the exhaust line. 

A pair of disposable flat disk filters is used to remove any particulates from the sample 

flow. The pressure drop across the filters is monitored by a Dwyer pressure gage 

attached to the flow line through 1.6 mm OD (1/16 in) stainless steel tubing. This small 

tubing minimizes any dead volume in these side lines. 

 

9

 The reason for dumping much of the flow before the instruments are reached is the need to dry and filter 

the instrument flows. The water traps or the filters would tend to clog rapidly if presented with the full 

flow. 
13

Sampled Exhaust

Excess Flow To Analyzers

Figure 5 Gas sampling system flow control panel. 

P

∆P

∆P

P

P

Manometer

Oxygen 

Meter

CO2 Meter

CO Meter

Helium

Detector

Instrument Rack Or

Ball Valve Press Regulator

Flow Control Valve Total Press Gauge

Flow Meter Diff Press Gauge

Mass Flow Controller Filter

Pump

P

∆P

Excess

Excess

Internal Bypass

Flow Balance

Flow Limit

Absolute Press in

Analyzer

1⁄4” to 1⁄2” Tube

1⁄4” to 1⁄2” Tube

1⁄4” Tube

1⁄4” Tube

1/16” Tube

1⁄2” Tube

1” Tube to

Bldg Vent

1” Tube to

Exhaust Hood

Via Manifold 3/8” Tube

From Heated

Sample Line

1⁄2” Tube

1/16” Tube

1⁄4” Coiled Tube

1⁄2” Tube

1⁄2” Tube

Zero and Span Gases

Excess Sample

Liquid Trap

1⁄2” Tube

DrieRite

Tube

Dry-Ice

Trap

Excess Zero/Span 

Gas to Exhaust

Sample Line

Blow Back Gas

Figure 6 Diagram of the gas analysis system for the 3 MW HRR Facility. 
14

The flow runs through a dry ice trap for moisture removal. The trap is filled with glass 

beads to improve heat and mass transfer, thus assuring that most of the water is removed. 

The trap is designed to easily handle a 30 min fire test without clogging. The pressure 

drop across the trap is monitored by a Dwyer pressure gage. The dry ice trap is backed 

up by a Drierite dessicant trap since a water aerosol (that may form at the trap inlet) could 

pass through the ice trap10. The flow emerging from the Drierite tube is filtered once 

again before going into the control room diaphragm pumps (GAST Model DOA-P109-

FB, rated at 54 L/min) to remove any Drierite dust. 

Flow from one of the control room pumps goes to the Servomex 540a oxygen analyzer; 

the flow from the other pump goes to the Siemens Ultramat 5e CO and CO2 analyzers. 

Each line has a manual valve, immediately downstream of the pump, to control the flow. 

The pressure gage in each outlet line helps monitor normal behavior; a faltering pump 

will be reflected in drift of its output pressure. 

All of the gas analyzers measure the partial pressure of the gas being sensed; thus they 

are all sensitive to total pressure variations in the instruments. Of particular importance 

is the need to keep the pressure in the instrument the same during a test as it was during 

the instrument calibration preceding the test. Since the oxygen measurement is the most 

critical, it is given extra attention in this regard. 

On the oxygen meter side of the pump, the flow is split just before a portion of it goes to 

the meter; the remaining flow is dumped to the exhaust line. The dumped flow passes 

through a pressure regulator (Porter Instrument Co., Model 8286) intended to help keep 

the flow in this path constant despite any upstream pressure variations. The flow rate is 

indicated by a Dwyer flowmeter downstream of the regulator. The flow to the oxygen 

meter is set by the controls on the oxygen meter itself. 

The Servomex 540a oxygen analyzer is a paramagnetic device that takes advantage of the 

strong magnetic susceptibility of oxygen molecules. The current output of the device is a 

direct measure of the oxygen partial pressure. Since the device measures partial pressure, 

it is sensitive to temperature as well as pressure. The sample cell is in a specially 

thermostatted enclosure to keep its temperature constant. The pressure in the sample cell 

depends on total ambient pressure and on the flow rate to the cell. As indicated above, 

special efforts are taken to hold the flow constant. The total cell pressure is actively 

measured (see below), and any change (from the level during the span gas measurement) 

is compensated in the data acquisition program. 

To assure further stability in the oxygen signal, an absolute pressure transducer (MKS 

Baratron Model 628, 0 to 133 kPa) is used to monitor the measurement cell pressure and 

to actively correct (in the data acquisition system) for any variations (with the value 

during calibration as the reference). A fitting internal to the oxygen meter, just upstream 

of the measurement cell, was modified to include a small diameter tube leading to this 

pressure transducer. 

 

10 The oxygen meter is sensitive to moisture in the gas sample. Past experience has shown that if about half 

of the Drierite tube length is showing a color change from blue to pink, the oxygen signal can begin to drift. 
15

The Siemens Ultramat 5e CO and CO2 analyzers both utilize a non-dispersive infrared 

absorption principle. Light from a broadband infrared source is split into two beams, one 

going through the sample cell and the other through a nitrogen-filled reference cell. A 

pair of double-layer detector cells is placed at the ends of both sample and reference cells. 

The four detector cells all contain a fixed percentage of the gas to be analyzed (CO or 

CO2). A chopper wheel causes a pulsating signal whose amplitude is proportional to the 

amount of CO or CO2 in the sample cell. 

The pump flow going to the CO and CO2 meters is in excess of instrument requirements 

so as to decrease the transit time and the excess is dumped as it enters the instrument rack 

area. The dump flow and the combined instrument flows are each controlled by a 

pressure regulator (Porter Instruments Model 8286). The flow to the two instruments is 

set using their built-in controls and flowmeters. 

2.2 Data Acquisition System 

2.2.1 Signal Input and Output 

The signals from the various measurement devices used for oxygen-depletion calorimetry 

determination of heat release rate are collected using the NIST Large Fire Research 

Facility data acquisition system (DAQ). The DAQ is currently configured to collect data 

both from instruments for calorimetry calculations and from other instruments or sensors 

being utilized for the experiment. 

The DAQ hardware components (manufactured by National Instruments) are stored in an 

instrument rack in the instrument control room of the Large Fire Research Facility. The 

DAQ components are separate from the computer and transfer of data between the DAQ 

and computer is via a PCI (Peripheral Component Interconnect) bus computer interface. 

The primary DAQ board has 16 analog input channels at 16 bit digital resolution and is 

capable of a sampling rate of 333 kHz. A total of 40 analog input signals are required for 

the calorimetry computation. Seven signal conditioning and multiplexing modules are 

employed and each module has a capacity for 32 input channels. The channels are 

scanned at a rate of 200 Hz and the signals are dumped to a single channel on the primary 

DAQ board. Each multiplexed channel is electronically averaged for 1 s and the digital 

value is stored to the computer. The signal conditioning features of the modules include 

an electronic filter (200 Hz low pass) and gains (up to 100 times). The gains are typically 

applied to low voltage signals such as those produced by thermocouples. Each module is 

connected to a shielded terminal block which collects signals up to 10 V amplitude and 

provides cold junction compensation for thermocouples voltages. 

2.2.2 Real Time Display 

The user interface for the DAQ is written in a LabView program to compute the heat 

release rate and display user-specified data and computations. The program is run on a 

Windows workstation with dual 733 MHz processors and is located in the instrument 
16

control room of the Large Fire Research Laboratory. The output is displayed on the 

workstation monitor and updated as the fire test occurs. 

During a fire test, all measurement channels are available for display. Calorimetry 

instrument readings, velocities, and heat release rate (both from the oxygen calorimetry 

calculations and the calibration burner) are displayed graphically with up to 6 min of 

history (see Figure 7). Additional plots of instrument raw voltages are available to assist 

with the zeroing and spanning processes. In addition, plots of other important, but noncalorimetry-related measurements are displayed when specified. 

Figure 7 Real time display of oxygen sensor cell pressure ratio (upper left), oxygen 

volume fraction (lower left) and heat release rate (right). 

Two marker channels are available and are used to keep track of specific events during a 

fire test. These markers are initiated with pushbuttons which are prominent on the 

display screen. Event marking puts simulated voltage spikes from 0 to 5 V and 10 to 15 

V for one second in the channels set up for the markers. Plotting the data in these 

channels along with a basic measurement or calculated value enables relating the events 

to the response of the calorimetry system or a particular component of it. Such events 

include the beginning of zeroing, spanning, and background averaging periods, ignition 

and extinguishment of the fire, and step increases and decreases that are of importance 

for delay and system smoothing characterization. 
17

2.2.3 Computation Algorithm 

Most data are stored as voltages without modification. (The exception is the temperature 

measurements, which are converted in real time to degrees Celsius.) A second data file is 

created simultaneously in which the raw voltages are converted to CO, CO2, and O2 dry 

volume fractions, pressures, etc. These are in turn combined to generate flow velocities, 

duct volume flow, duct mass flow, and heat release rate. For heat release calorimetry 

testing, data are combined after estimating the relative delays for each instrument 

(described below) and synchronizing the data generated by a given moment of the fire. 

Each channel of raw or computed data has an associated delay time. The delay time is 

the period between when a fire event occurs under the hood and when the particular 

sensor begins to respond to the change associated with the fire event. There are two main 

categories of delay. The first is for those sensors located at the sampling section of the 

duct which include pressure transducers and thermocouples. These sensors respond in 

less than a second to changes in their environment, but the input from an event under the 

hood is delayed in reaching the sensors by the flow time for the gas at the flame tip to 

enter the hood and progress through the duct to the sampling section. The second 

category of delay is for the gas analyzers located in the experiment control room. These 

instruments experience the duct flow delay as well as a sampling line flow delay and their 

own inherent sample processing delays (response times). These delays, which vary with 

the exhaust flow, are determined from specific tests that investigate instrument response 

to step changes under the hood. 

The data acquisition and analysis program uses the delays to combine the outputs of 

sensors and instruments in such a way that the responses to particular phenomena under 

the hood are synchronized. In other words, output from fast responding sensors is stored 

in memory to be combined with output from the more slowly responding sensors at the 

appropriate time. Thus the faster sensor outputs are actually delayed in their 

combinations and calculations while the slowest outputs are utilized instantly. 

The concept of storing varying histories of sensors with different response times is 

straightforward. The maximum delay time experienced by any sensor in the whole set of 

sensors defines how much history is stored for all sensors in the “history array”. A delay 

time for each sensor is converted to an integer number of seconds. This also represents 

the position in the array that corresponds to the appropriate age of the data relative to that 

of the slowest sensor (the age and position correspond directly due to the 1 Hz sampling 

rate). 

As the fire test proceeds, the calculations are grouped in several sets. Each set builds on 

variables calculated in the previous sets. The algorithm flowchart in Figure 8 displays 

these calculation steps. 
18

Initialize files, headings, most arrays. 

Initialize analog input process, more 

arrays, pulse counting. 

Initialize main loop for scanning voltages 

and computations. Check flow setting and 

establish delays. 

Convert raw voltages to physical values, 

generating, e.g., pressures, temperatures. 

Combine or run further calculations on 1st

level converted values, generating, e.g., 

gas concentrations, velocities. 

Collect raw voltages for 

“raw” file. Scan for event 

marker inputs. Average 

voltages for zeroes, spans, 

background when 

activated. 

Combine or run further calculations on 1st

and 2nd level converted values generating 

velocity averages and oxygen depletion 

factor. 

Calculate calorimetry HRR based on 

previously converted values. 

Extract latest time slice of synchronized 

data and combine in time-adjusted array. 

Write “raw” and “adj” data to 

files. 

Display charts of HRR, 

velocity profile, etc. 

Store latest data in history file as 

long as slowest sensor delay. 

Combine or run further calculations on 1st

to 3rd level converted values generating 

volume and mass flows in exhaust duct. 

Store latest data in history file as 

long as slowest sensor delay. 

Store latest data in history file as 

long as slowest sensor delay. 

Store latest data in history file as 

long as slowest sensor delay. 

1 

Yes

No

Start loop 

again? 

End

Write “zs” zero and span file. 

Close all files. 

1

Figure 8 Heat release rate computation algorithm. 

In the first set, the calculations include all temperatures, basic O2 volume fraction, CO 

and CO2 range factors, pressures, and mass losses. The second set of calculations 

includes those for oxygen volume fraction pressure correction, CO and CO2 volume 

fractions, bi-directional probe velocities, water vapor volume fraction at ambient 

temperature, and average duct gas temperature near the sampling section (and the 

uncorrected natural gas flow and pressure for calibration burns). The third set of 

calculations includes the average duct velocity, and the oxygen depletion factor, φ (and 

the heat release rate of the natural gas for calibration burns). The fourth set consists of 

calculations of the volumetric duct flow and the mass flow in the duct. The last step 

calculates the heat release rate based on the duct mass flow, φ , and the O2, CO, CO2, and 

H2O volume fractions. 
19

2.2.4 Data Storage 

Three files are generated during a fire test with suffixes “raw,” “adj,” and “ZS:” The 

“raw” file records the raw voltage values, thermocouple temperature values, and the two 

event marker channel values. 

The “adj” file records the simple one-step conversions applied to the raw voltages to give 

values in terms of scientific significance such as species volume fraction and pressure. In 

addition the results of calculations for velocity, mass flow rate, and oxygen depletion 

factor are stored. Finally, the computation of the heat release rate is stored. The “raw” 

and “adj” files are updated each second with each sample of the data acquisition system 

and the associated time stamps are stored. 

The “ZS” file lists the voltages generated by the zeroing and spanning processes for the 

gas analyzers as well as the calibration signals for the response of pressure transducers, 

load cells, etc. For some of the instruments, such as the gas analyzers, the calibration 

curve is linear and interpolation between two calibration points enables assignment of 

volume fractions to the recorded voltages. Other instruments, such as the pressure 

transducers and load cells, are generally non-linear or nearly linear and have been precalibrated over the necessary range. These calibration equations require a single point 

each to enable conversions from voltage to, e.g., Pa or kg. 

2.3 Calibration Burner System 

Because the calculation of heat release rate requires a large number of individual 

measurements, it is essential to have an independent confirmation of measurement 

accuracy. This is accomplished by burning an accurately measured flow of natural gas 

having a well-defined heat of combustion. 

Figure 9 Photo of 3 MW fire from the natural gas burner. 
20

The natural gas burner, shown in Figure 9, has been designed, fabricated and tested for 

this purpose. The burner can produce flows that correspond to a wide range of heat 

release rates, from 50 kW to over 6 MW in any increments specified by the user. The 

burner and flow system are configured to use natural gas as supplied by the local gas 

company. The energy content of the gas supply is measured by a natural gas calorimeter 

(COSA CW95) located in the instrument control room. Operation of the burner is 

automated with control and monitoring executed from the DAQ system. 

2.3.1 Burner 

The natural gas burner has a height of 96 cm (measured from floor to exit holes) and plan 

view dimensions of 112 cm x 150 cm (width x length). It is capable of producing 

controlled fires up to 6 MW using natural gas as fuel. The natural gas is delivered to the 

combustion zone above the burner through eleven primary rows of tubes. Two pilot 

tubes are located below and perpendicular to the primary tubes. The eleven primary 

tubes are grouped into four separate banks, which are independently opened and closed 

using solenoid valves. Operation of the four primary banks is dependent upon the 

amount of energy released in order to regulate flame-height, maximize efficiency and 

minimize the risk of flame blow-off. Each bank consists of three tubes except for Bank 1 

which consists of two tubes. Each primary tube is split in half and capped at the end to 

allow for thermal expansion, thereby minimizing stress on the burner. Each tube has 56, 

3 mm diameter holes drilled every 2.54 cm. Flow is initiated to the pilot tubes by a 

solenoid valve and the pilots are ignited using a 10,000 V spark igniter. Figure 10 is a 

photograph of the burner. 

Figure 10 Calibration burner. 

Pilot Tubes 

Primary Tubes 

10 cm Distribution Pipe

Solenoid Valves 

(3 of 5) 

Pilot Tube 

Pressure 

Regulator 
21

The burner is supplied with natural gas by a 7.5 cm diameter (i.d.) flexible stainless steel 

wound hose line, which feeds into the burner through a 10 cm diameter (i.d.) main 

distribution pipe. The main distribution pipe feeds both the two pilot lines and the 11 

primary tubes. An electronic photoelectric sensor monitors the pilot flame and is linked 

to a shutoff valve to assure that a flame is burning when the gaseous fuel is flowing. 

2.3.2 Flow Control and Flow Measurement 

The burner is supplied with natural gas through an overhead gas train, which controls and 

measures the amount of gas supplied to the burner. Gas pressure is regulated to 

approximately 100 kPa above atmospheric pressure prior to reaching the overhead train. 

The gas train consists of a filter, a flow control valve, a volumetric gas flow meter, and a 

safety shut-off valve. The flow control valve is a pneumatically driven, programmable, 

high precision valve that is controlled by current provided by the data acquisition system. 

The volumetric flow meter is a rotary displacement meter that provides an amplified 

pulse output that is detected by the data acquisition system. A thermocouple and a 

pressure transducer are located in the volumetric flow meter so that mass flow may be 

deduced from the volumetric flow measurement. Figure 11 is a photograph of the gas 

train installed overhead in the Large Fire Research Laboratory. 

\

1

2 3

4

Figure 11 Overhead gas train including the (1) control valve, (2) rotary 

displacement meter, and (3) emergency shut-off valve. The exhaust duct (4) is also 

shown. 

Figure 12 is a schematic diagram of the burner and the systems used to measure and 

control the flow of gas. The automated safety system is also shown on the diagram and 

includes a photoelectric sensor (FireEye), the Safety Panel, and Emergency Cut-Off 

switches. The input and outputs to the Safety Control panel are shown schematically in 

Figure 12. The panel is used to initiate the gas flow and to cut-off the gas flow in the 

event of an emergency. The manual cut-off closes the emergency shut-off valve labeled 

as item (3) in Figure 11. For emergencies, manual cut-offs are also positioned at doors 

on the east and west sides of the Large Fire Research Laboratory. 
22

90 O ELBOW WITH 

3/8 IN. NPT PORT 

FOR N2 PURGE 

INJECTION

FIRE EYE (UV)

IGNITERS

PC/

LabVIEW

SHUTOFF VALVE 

WITH MANUAL 

OVERRIDE

MANUAL ON/OFF 

2 1⁄2” PIPE

DISPLACEMENT 

METER

PRESSURE 

REGULATOR

MASON EILAN 

FLOW CONTROL 

VALVE

MANUAL ON/OFF

90

3”-DIA. STEEL BRAIDED 

FLEXIBLE HOSE

PILOT BURNERS

1.02 m 0.51 m

ON/OFF SOLENOID 

ARRAY

3/8 ” PIPE

2 1⁄2” PIPE

3” HOSE

4” PIPE

PRESSURE

TEMPERATURE

Safety 

Panel

LEGEND

VOLUMETRIC 

FLOW (PULSES)

90 O ELBOW WITH 

3/8 IN. NPT PORT 

FOR N2 PURGE 

INJECTION

FIRE EYE (UV)

IGNITERS

PC/

LabVIEW

SHUTOFF VALVE 

WITH MANUAL 

OVERRIDE

MANUAL ON/OFF 

2 1⁄2” PIPE

DISPLACEMENT 

METER

PRESSURE 

REGULATOR

MASON EILAN 

FLOW CONTROL 

VALVE

MANUAL ON/OFF

90

3”-DIA. STEEL BRAIDED 

FLEXIBLE HOSE

PILOT BURNERS

1.02 m 0.51 m

ON/OFF SOLENOID 

ARRAY

3/8 ” PIPE

2 1⁄2” PIPE

3” HOSE

4” PIPE

PRESSURE

TEMPERATURE

Safety 

Panel

LEGEND

VOLUMETRIC 

FLOW (PULSES)

Figure 12 Schematic diagram of the natural gas burner system. 

2.4 Design Considerations 

2.4.1 System Time Delays and Time Response 

The computed heat release rate is displayed on the computer monitor during a test; 

however, the numbers lag behind the actual fire output. There are several components to 

the overall transient response of the system to an instantaneous event in the fire being 

measured. Initially there will be no response since the information about a change has 

not reached the analyzers whose data go into the heat release rate calculation. So the first 

component of transient response is the gas travel time from the flame tips to the location 

in the duct where the gases are sampled; this varies with the hood flow rate. Next there is 

the gas travel time from the gas sampler on the roof to the inlet of the instruments in the 

control room (which includes the travel time through the plumbing in the control room). 

These are independent of hood flow rate but depend on the pumping speed of the system 

and the use of bypassed flows. Finally there is the response time of the instruments 

themselves. This is fixed when the flow rate to each instrument is set at its manufacturerrecommended value. Along the entire path, the gas does not travel as a rigid slug; rather 

it tends to undergo mixing with gases originating later as a result of various turbulent 

mixing processes. This process, called dispersion, also affects time response of the 

overall system to a change in the fire behavior. Careful design helps minimize this effect. 

In general one needs to measure the lag time (total sample flow time) from the fire to the 

instrument inlet. This has a differing value for the oxygen meter and the CO/CO2 meters 

since they involve differing flow paths (and flow rates) in the control room. Lag time is 

used to synchronize the measurements that go into the heat release rate calculation. One 

also needs to measure and take note of the system response times for the individual gases. 

These are to be kept within a certain range since they affect how closely the system will 
23

follow a given time-varying heat release rate process from the fire being measured. In 

particular, system response times affect accurate measurement of brief heat release rate 

peaks (see section on transient effects below). 

As a first approximation, a few of these times break out as follows: 

• Duct flow time (fire to sampling point at 9.9 m3

/s) ≈ 7 s 

• Gas sample flow time from roof to control room ≈ 2 s to 3 s 

• Instrument response times at nominal flow rates ≈ 5 s to 10 s 

The total system delays are somewhat longer than the sum of these numbers, however; 

for example, the total delay is 25 s for a particular experiment discussed below. 

Precise values of the lag time for each instrument are needed since they must be 

incorporated into the data handling that results in a correct calculation of heat release 

rate. Measurement of these times is facilitated by a heat source that can be changed 

abruptly in magnitude (up and down). This capability is achieved via the data acquisition 

computer which commands the degree of opening of the gas flow control valve in the line 

to the multi-tube calibration burner; this system itself has an approximately one second 

lag time in response to computer commands. 

The first time interval, the flow delay to reach the gas sampling plane in the duct on the 

roof, is measured by putting an event mark in the test file when the gas flow to the burner 

is turned on and then by finding the time at which any11 of the thermocouples on the bidirectional probes begins to register a gas temperature increase. Reference [15] 

recommends a specific value of a temperature increase of 3 K12 and that is what is used 

for the NIST system. Again, this time interval varies with the hood flow rate. 

The second component of the lag time for a given instrument, the time to go from the gas 

sampler on the roof to the inlet of the instrument, is a constant if the speeds of the roof 

and control room pumps are constant and the fraction of the excess sample flow dumped 

at the first flow split is unchanged. However, as noted, it can vary between the oxygen 

meter and the CO/CO2 meters, because the flow paths differ from the laboratory pump(s) 

onward. This component of the lag time is found by noting the time between the above 

thermocouple response and the first instrument response to the fire-altered gases. 

Reference [15] recommends taking the time when the oxygen meter reading shows an 

increment of 0.05 % and the CO2 meter an increment of 0.02 %; these are also adopted 

for the NIST system. Note that the CO level from a gas burner can be difficult to read. 

One can safely assume that the lag time component for it is the same as for the CO2 meter 

since they are effectively in the same flow system. 

 

11 The duct flow has been found to be well-mixed at flow rates of 9.4 m3

/s and above. It is prudent to 

check more than one thermocouple at lower flow rates. 

12 To get a sharply readable increase of this size may require a fire of several hundred kW. Larger hood 

flows require a larger fire. Too small a fire will give erratic results. 
24

The system response time is defined in Ref. [15] as the time for the oxygen meter13 to go 

from 10 % to 90 % of the change induced by a step change in fire size. Again, any and 

all dispersion in the system, from the fire plume tip all the way to the gas analysis 

instrument, contributes to the time it takes for this signal to change from 10 % to 90 %. 

This dispersion smears the original time sequence of events in the fire. In particular, it 

decreases the reading of any heat release rate peak that occurs on a time scale comparable 

to or shorter than this 10 % to 90 % response time of the system. This raises an issue that 

is discussed further in the next section. 

2.4.2 Transient Effects on Peak Heat Release Rate 

Fire regulations are frequently based on maximum allowed rates of heat release. 

Different calorimeter facilities have differing time responses and thus differing 

tendencies to smooth out a brief peak. Some facilities use an open hood while others use 

a compartment which allows smoke to build up prior to spilling into the hood, thus 

adding a delay in measurement of heat release rate. 

Messerschmidt and Van Hees [15] discuss this transient response issue in the context of 

the European Community’s Single Burning Item Test, which is based on heat release rate 

measurement. It is shown that different laboratories could end up rating a product 

differently in this test solely as a result of the differing transient behavior characteristics 

of their measurement systems. They recommend that the oxygen meter’s apparent 

response time (defined as above) fall within a window of 9 s to 12 s, in order to minimize 

the tendency of differing labs to rate a product differently in the Single Burning Item test. 

Another aspect of transient system response is related to noise in the calculated heat 

release rate. System noise (a result of the random variations in all of the variables 

entering into the heat release rate calculation) can affect the apparent magnitude of a heat 

release peak, and thus its uncertainty. The effect depends on the root mean square noise 

level and its frequency characteristics relative to the peak duration. An approach that is 

sometimes used is to judge a tested product’s performance on the basis of a heat release 

rate curve subjected to a running time averaging process. The longer the time interval 

used for averaging, the more the noise is filtered out but also the more real, rapid heat 

release rate changes are damped. This raises the question as to how rapid a heat release 

rate fluctuation matters when an object is burning in some context such as a compartment 

fire. There may not be a unique answer to this and it is a subject of planned research at 

NIST. For now one can estimate that in a compartment fire context, upper layer gas 

temperature changes and second object ignition processes begin to respond significantly 

to heat release rate alterations in a time of the order of 5 s to 15 s, suggesting this range as 

the appropriate upper limit for time smoothing of the data. In practice, it is easy to use 

graphing software to examine the effect of varied time smoothing intervals on the heat 

release rate peak of a test object. For mattress fires it was found that 20 s smoothing 

reduced the apparent peak values by a couple of percent. In Ref. [15], a 30 s interval was 

 

13 The CO and CO2 meters’ response times (inherent to the instruments, not the overall system) tend to be a 

little faster than the oxygen meter and thus do not limit the system response. In any event, they tend to be 

secondary in importance to HRR calculation. 
25

used without any discussion. The smoothing interval, if any, should be reported along 

with the heat release rate data. 

Determining the Role of Transient Effects on Measured Peak Heat Release Rate and 

Energy Conservation

As part of the present work, a series of experiments was performed to characterize the 

response of the 3 MW heat release rate calorimeter to square-wave pulses in terms of 

peak heat release rate, the width of the peak, and the conservation of energy. The results 

of these experiments were used to establish a criterion for assessing whether the peak 

heat release rate for an experiment is being underestimated because of the time response 

of the system. 

In principle this is a problem which can be treated as a deconvolution of the actual heat 

release rate vs time from the measured heat release and the system response 

characteristics [16,17]. Thus one solves an integral equation relating observed heat 

release rate to the time integral of actual heat release rate modulated by the measurement 

system response. Differing mathematical techniques have been demonstrated in the 

noted references. The inferred correction to the apparent heat release rate history is 

sometimes large (O(100%)), particularly for cases in which the heat release rate peak 

width is comparable to the time response of the measurement system. However, it is 

pointed out in Ref. [18] that the deconvolution process is quite sensitive to noise in the 

data and to the errors in the system response function. Thus it is of unclear value here 

where both of these issues are significant. Rather than proceed in this manner, we focus 

on characterizing the time response domain in which the 3 MW hood system gives 

accurate results. Deconvolution remains an area for further exploration. 

The basic approach to assessing system time response was to produce a square-wave 

pulse heat release rate using the natural gas calibration burner described in Sec. 2.3. The 

peak heat release rate was set at nominal values of 0.5 MW, 1.0 MW, and 2.5 MW using 

a flow control valve. A signal was manually initiated from the computer control panel to 

open the flow control valve to a preset value and a second signal was sent to reset the 

flow to the pilot level. The loop time of the program for controlling/acquiring data is 1 s. 

The heat release rate of the burner is computed from the heating value of the natural gas, 

its pressure, and the flow rate. As discussed in Sec. 2.3.2, the flow meter is a rotary 

device which displaces a fixed volume of gas four times per revolution. Digital pulses 

are produced at a rate of about 1 pulse per 2 L of gas. A heat release rate of 1.0 MW 

corresponds to about 7 pulses per second and 2.5 MW to about 18 pulses per second. 

The relatively low number of pulses per second limits the resolution for 1 Hz 

measurements such as needed in this study, but it does allow an accurate estimate of the 

average heat release rate. 

A Schmidt-Boelter type radiometer with a time constant of 0.1 s to 0.2 s was used to 

provide a more rapid response to the actual peak heat release rate. The radiometer was 

set up a distance of 4 m from the burner with the view including the entire flame volume. 

The corrected burner peak heat release was obtained by multiplying the radiometer output 
26

times a calibration factor determined for each of the three heat release rates. To 

determine this calibration factor, the heat release rate computed from the calibration 

burner volume flow measurement and radiometer output were averaged over a 20 s 

period during the steady burning for each of the three heat release rates. The calibration 

factor is computed as the ratio of the average heat release rate to the average radiometer 

voltage. The ratios have a dependence on fire size with values decreasing by 30 % for a 

five-fold increase in the heat release rate (2.62 GW/V at 0.5 MW, 2.20 GW/V at 1 MW, 

and 1.85 GW/V at 2.5 MW). 

Figure 13 provides a detailed comparison for the burner heat release rate based on the 

flow meter and the output from the radiometer. Qualitatively the pulse shapes are similar 

as measured by both devices. The limited signal resolution of the flow meter is apparent 

from the discrete change in heat release rate around 4955 s and again around 4970 s. 

There appears to be a one to two second delay in the processing of the pulse counting 

relative to the radiometer measurement. This delay was observed in all the tests. The 

inflection point of the radiometer output near the end of the pulse is thought to result 

from continued flow through the burner after the valve for the flow controller has been 

reset to the baseline level. The gas continues to drain as pressure in the gas line to the 

burner equilibrates to the lower baseline level. 

0

0.3

0.6

0.9

1.2

-0.1

0

0.1

0.2

0.3

0.4

0.5

4950 4955 4960 4965 4970 4975

NG burner, MW

Radiometer, mV

Time, s

Figure 13 The square-wave pulse from a 15 s release of natural gas for a 1 MW heat 

release rate as measured by a radiometer and the burner flow meter. 

The measurement sequence began by setting the calibration burner at a heat release rate 

of 0.10 MW. This pilot level of heating is needed to efficiently ignite the rapidly 

increasing gas flow. The mass flow controller was set for a flow corresponding to 1 MW 

heat release rate and then a switch was activated which in turn opened the mass flow 
27

controller valve. After a fixed time, the sequence was reversed and the heat release set 

back at 0.1 MW. After the calorimeter reached a steady heat release at this low value, a 

repeat run would be made following the same procedure as above. Typically three runs 

were made for each of four pulse times: 30 s, 15 s, 10 s, and 5 s. At least one 

measurement was also made for a 60 s pulse. This sequence was then repeated for heat 

release rates of 0.5 MW and 2.5 MW. 

0

0.2

0.4

0.6

0.8

1

1.2

-0.1

0

0.1

0.2

0.3

0.4

0.5

4150 4200 4250 4300

HRR, MW

Radiometer, MV

Radiometer Calorimeter

NG burner 30 s pulse

15 s pulse

0

0.2

0.4

0.6

0.8

1

1.2

-0.1

0

0.1

0.2

0.3

0.4

0.5

5020 5040 5060 5080 5100 5120 5140

HRR, MW

Radiometer, MV

10 s pulse

5 s pulse

Time, s

Figure 14 The measured output versus time for the NG burner (), calorimeter (○), 

and radiometer (solid line) resulting from 1 MW square-wave pulses of 30 s, 15 s, 10 

s, and 5 s. 

Figure 14 illustrates the characteristics of the square-wave pulse shapes for the 1 MW fire 

for the four different pulse widths. The 25 s time delay for the calorimeter is a result of 

the gas sampling time together with the nominal 9 s response time of the oxygen 

analyzer. The relatively low resolution of the natural gas flow measurement is apparent 

from the step changes in the heat release rate. The radiometer provides a better 

representation of the heat release pulse shape since it responds to the flame intensity, and 

it has a higher resolution (and faster response time) than the gas flow meter. For the 30 s 

and 15 s pulses, the peak in the calorimeter output is close to that of the burner and the 
28

full time widths at the half peak height (FWHH) are similar though the curve shape 

produced by the calorimeter is rounded rather than square. The calorimeter peak is 

slightly reduced for the 10 s pulse and is reduced by about a factor of two for the 5 s 

pulse. In these two cases the FWHH values are also increased relative to the radiometer 

output. 

0

0.2

0.4

0.6

0.8

1

1.2

-0.1

0

0.1

0.2

0.3

0.4

0.5

4950 4960 4970 4980 4990 5000 5010 5020

Calorimeter, MW

Radiometer, mV

NG burner peak HRR

0

0.5

1

1.5

2

2.5

3

0

0.3

0.6

0.9

1.2

1.5

7400 7410 7420 7430 7440 7450 7460 7470

Calorimeter, MW

Radiometer, mV

Time, s

NG burner peak HRR

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

-0.05

0

0.05

0.1

0.15

0.2

0.25

5800 5810 5820 5830 5840 5850 5860 5870

Calorimeter, MW

Radiometer, mV

NG burner peak HRR

Figure 15 The heat release rate and the radiometer output as plotted versus time 

for 15 s pulses for HRR of 0.5 MW, 1 MW, and 2.5 MW 
29

The pulse experiments for the other two heat release rates provided similar results. It is 

seen in Figure 15 for the 15 s square-wave pulses for the three heat release rates that the 

peak heights are similar to those of the NG burner, the FWHH are similar to those from 

the radiometer, and that the scaled shapes of all three calorimeter-generated curves are 

similar. There are slight differences in the radiometer-generated pulse shapes such as a 

flatter top section and a more abrupt decrease in the HRR at the end of the pulse for the 

larger heat release rates. These subtle changes do not seem to be reflected in the 

calorimeter-generated curves. 

The total heat released by the pulse is equal to the area under the burner output curve. 

For energy conservation for the calorimeter system, the area under the calorimeter curve 

should equal the area under the burner curve. In Figure 14 the two areas appear to 

approximately match for the longer two pulses, but it is difficult to determine if this is the 

case for the shortest two pulses. A quantitative assessment of energy conservation was 

made by determining the total integrated heat released computed based on the burner 

flow and on the calorimeter output. The integrated area is that above the nominal 100 

kW baseline. The baseline correction is estimated as the average of the baseline values 

before and after the pulse. The variation in the average ratio of the integrated calorimeter 

output to the integrated burner output for a fixed heat release rate (see Table 1) is within a 

range of 10 % except for the 1 MW fire, for which the ratio decreased by as much as 11 

% for the 5 s pulse. It is thought that the low number of pulses, about 3, counted per 

second is responsible for this decrease. 

It was decided to again take advantage of the more rapid response time of the radiometer. 

A corrected total burner heat release was obtained by integrating the radiometer output 

and then multiplying by a calibration factor determined for each of the three heat release 

rates. The calibration factor was selected so that the radiometer “energy” equaled the 

integrated heat release rate from the gas burner for the 30 second experiment. The flow 

meter accuracy of the gas burner is improved about five-fold with the longer integration 

time. Using this approach, it was found that the average ratio of energies deviates by 6 % 

or less for the shortest to the longest pulse for any one burner setting. There is an 

increased uncertainty for the shortest pulse length. For the 5 second pulses, the range in 

the standard deviation is 4 % to 7 % of the mean compared to a more typical range of 1 % 

to 3 % of the mean for the longer pulses. 

Energy conservation has been established to within the resolution of the experiment (see 

Table 1). This basic premise for obtaining a quantitative measurement of the total energy 

released regardless of how small a pulse is that the total oxygen consumed in the 

combustion reactions is obtained from the integrated oxygen analyzer output. This 

premise has been validated. 
30

Table 1 Energy conservation results 

Nominal pulse 

width, s Integrated burner output

Integrated calorimeter output

Integrated radiometer output

Integrated calorimeter output a

0.5 MW

30 0.97 ± 0.01 0.97 ± 0.04

15 1.01 ± 0.03 0.98 ± 0.02

10 0.96 ± 0.01 0.92 ± 0.03

5 0.90 ± 0.06 0.92 ± 0.04

1 MW

30 1.06 ± 0.02 1.06 ± 0.02

15 0.96 ± 0.03 1.05 ± 0.03

10 0.98 ± 0.02 1.10 ± 0.02

5 0.89 ± 0.04 1.05 ±0.06

2.5 MW

30 1.02 ± 0.01 1.02 ± 0.01

15 1.02 ± 0.02 1.05 ± 0.01

10 1.02 ± 0.02 1.07 ± 0.01

5 0.97 ± 0.04 1.06 ± 0.07 a

 The integrated radiometer output for each of the three heat releases was multiplied by a factor to match 

the total burner heat release rate for the 30 s square-wave pulses. 

A summary of the average peak HRR and FWHH and the associated standard deviations 

are given in Table 2 and Table 3 for each set of three or more repeat measurements. The 

FWHH is computed based on the half-height above the baseline value. The peak HRR is 

not adjusted for the baseline value. For pulse lengths of 15 s and longer, the calorimeter 

peak is within 9 % of the burner peak. The peaks correspond to the instantaneous 

maximum values recorded by the calorimeter, the corrected burner, or the radiometer. 

The five second-smoothed peaks were found to yield similar results except for the 

shortest pulse widths. The results indicate that the calorimeter peak is about 15 % 

smaller than the burner peak for the 10 s pulse and about 50 % smaller for the 5 s pulse. 

The results in Table 2 indicate that the calorimeter value on average is a factor of 1.03 ±

0.04 greater than the corrected burner peak for square-wave pulses having widths equal to 

or greater than 13 s. The calorimeter significantly underestimates the peak height for 

square wave pulses lasting less than 10 s. 

The corrected calibration peak heat release rate is not affected by the pulse width down to 

5 s. The standard deviation of the burner peak heat release rate for the four pulse widths 

is at most 5 % of the average for any of the three HRR (see Table 3). 

The FWHH time based on the calorimeter for the 1 MW fire is at most 10 % longer than 

the radiometer time for the 15 s and 30 s pulse experiments while it is about 25 % longer 

for the 10 s pulse and more than a factor of two longer for the 5 s pulse. Another 

observation is that the FWHH measured by the calorimeter remains relatively constant as 

the pulse width is reduced from 10 s to 5 s. In this limit, the calorimeter response is 

controlled by the time response of the oxygen analyzer so that the response approaches a 

constant independent of how short the square-wave pulse might be. 
31

Table 2 Gas burner square-wave pulse results for peak HRR and FWHH 

Pulse width, 

s Burner peak

Calorimeter peak Radiometer 

FWHH, s 

Calorimeter 

FWHH, s 

0.5 MW 

60a 1.03 59 58 

30 0.97 ± 0.02 28.0 ± 0.2 29.3 ± 0.3 

15 1.01 ± 0.06 12.6 ± 0.1 14.4 ±0.2 

10 0.89 ± 0.05 8.7 ± 0.6 11.1 ± 0.4 

5 0.58 ± 0.02 3.8 ± 0.6 8.9 ± 0.7 

1 MW 

60 1.06 ± 0.03 58.6 ± 0.7 59.3 ± 0.8 

30 1.06 ± 0.01 29.3 ± 1.2 30.0 ± 1.7 

15 0.96 ± 0.01 13.8 ± 0.5 15.2 ± 0.7 

10 0.82 ± 0.06 8.1 ± 0.5 10.8 ± 0.3 

5 0.52 ± 0.03 3.4 ± 0.2 8.9 ± 0.3 

2.5 MW 

60a 1.09 58 58 

30 1.05 ± 0.03 28.2 ± 0.9 28.2 ± 0.8 

15 1.04 ± 0.02 14.4 ± 0.1 14.7 ± 0.2 

10 0.88 ± 0.03 8.7 ± 0.8 10.6 ± 0.1 

5 0.47 ± 0.02 3.4 ± 0.2 10.3 ± 0.4 

 aOnly one test performed. 

The response of the calorimeter to the square-wave pulses suggests a possible method for 

checking whether the peak heat release rate is changing too rapidly to obtain an accurate 

heat release rate with the calorimeter. If the FWHH is 15 s or longer for the HRR 

calorimeter, then one has confidence from this study that the heat release is accurately 

represented by the data. However, if the FWHH is less than 11 s the peak heat release 

will likely be underestimated by 15 % or more. It does not appear to be possible to 

correct the peak heat release rate based on the FWHH, since small changes in the FWHH 

can result in a factor of two or more increase in the peak heat release rate, as seen in the 

data in Table 2. On the other hand, the square-wave pulse data may be useful in testing 

algorithms such as that developed by Abramowitz and Lyons [17] for deconvoluting the 

actual heat release rate from the measured values. 

A second important issue is the effect of the random system noise on the apparent 

magnitude of the heat release peak and thus the uncertainty in the peak height. The 

random noise sources include the gas analyzers, the pressure transducers for measuring 

velocity, the thermocouples, real variations in the temperature, velocity, and gas levels 

being measured, variations in the pumps and exhaust fans, and changing ambient 

humidity and temperature. By looking at the peak heat release rate in repeat tests under 

nominally identical conditions, we can assess the effect of the calorimeter related noise 

sources on the measured peak. A convenient measure of repeatability is the coefficient of 

variation (CV), which is defined as the ratio of the standard deviation for repeat 
32

measurements to the average value. We find that the CV for corrected peak burner heat 

release rate (see Table 3) is at most 0.05 for any of the 12 set of tests with at least three 

repeats each and is 0.02 or less for 5 of the 6 sets of tests with pulse widths of 15 s or 30 

s. The CV for the calorimeter is also found to be in the range of 0.01 to 0.02 for 5 of the 

6 sets of tests with pulse widths of 15 s or 30 s. The near equality of the CV’s suggest 

that the calorimetry repeatability may be limited by the slight lack of repeatability in the 

square-wave pulse produced by the burner Larger CVs up to 0.10 are obtained for heat 

release rate by the calorimeter for the 5 s and 10 s square-wave pulses. 

Table 3 Gas burner square-wave pulse coefficient of variation (CV) 

Nom. pulse 

width, s 

Corrected burner 

peak, MW 

Burner CV Calorimeter 

peak MW 

Calorimeter CV 

0.5 MW 

30 0.59 0.01 0.56 0.02 

15 0.56 0.04 0.56 0.03 

10 0.56 0.05 0.45 0.02 

5 0.54 0.04 0.34 0.10 

Avg ± σ 0.56 ± 0.02 

1.0 MW 

30 1.00 0.02 1.06 0.02 

15 1.05 0.01 1.00 0.01 

10 1.05 0.03 0.85 0.06 

5 1.04 0.01 0.53 0.04 

Avg ± σ 1.04 ± 0.02 

2.5 MW 

30 2.53 0.02 2.63 0.02 

15 2.47 0.01 2.57 0.01 

10 2.52 0.01 2.23 0.04 

5 2.46 0.02 1.14 0.02 

Avg ± σ 2.50 ± 0.04 

It is encouraging that the contribution of the system noise to CV for the peak heat release 

rate for the 30 s and 15 s square-wave pulses is typically 0.01 to 0.02. However, this 

does not mean that the CV for repeat burns of identical furnishing items, consumer 

products, or building materials will be as small. The CV in HRR measurements for such 

materials may be far greater than 0.02 because the identical items may not really be 

“identical” due to material and/or construction variation or because the flame spread 

process occurring on real objects is very sensitive to the details of the ignition process as 

well as the ambient flow and temperature field. Fire growth processes on real objects of 

any complexity (e.g., mattresses or items of upholstered furniture) are notoriously 

difficult to reproduce precisely. Since these help dictate the shape of the heat release rate 

curve from the overall object, and its peak, one can expect appreciable variations in 

measured heat release rate for such items, even with an ideal calorimeter. 
33

-0.1

0

0.1

0.2

0.3

0.4

0.5

-0.1

0

0.1

0.2

0.3

0.4

0.5

4940 4950 4960 4970 4980 4990 5000 5010

Radiometer, mV

Gaussian, mV

Time, s

8 s smooth

Figure 16 Comparison of the effect of 8 s smoothing on the peak height for a 

square-wave pulse vs. a Gaussian pulse. The peak heights for the square-wave pulse 

and Gaussian are 0.488 mV relative to the baseline, the FWHH is 13 s, and the 

standard deviation of the Gaussian is 5.52 s. 

The analysis presented so far has computed measured time response characteristics for 

square-wave pulses. In reality it is likely that the peak will be more rounded in shape. A 

Gaussian peak is a useful example. In the domain where calorimeter response affects the 

results, one might expect the calorimeter peak heat release rate for a square-wave to be 

larger than for a Gaussian with the same FWHH. The estimate here is based on the 15 s 

square-wave pulse for the 1 MW fire plotted in Figure 13 with a FWHH of 13.0 s. A 

Gaussian pulse was modeled with a standard deviation of 5.44 s corresponding to the 

13.0 s FWHH value. The magnitude of the Gaussian peak above the baseline was taken 

to be equal to the corresponding measured value. To simulate the system response of the 

calorimeter, an eight second smoothing was used for both the “square-wave” pulse and 

the Gaussian. As indicated in Figure 16, the Gaussian peak is decreased by about 10 % 

by including the time response of the system through smoothing while the decrease for 

the Square-wave peak is about 3 %. The reason for the closer approach to the peak for 

the square-wave is the ten-second duration at the maximum value compared to a onesecond duration for the Gaussian. This effect offsets the much more rapid rise for the 

square-wave compared to the Gaussian. The difference in the peak heights for the two 

pulses will decrease for wider pulses and increase for shorter pulses. 

The ability to reproducibly generate nominal square-wave heat pulses has been important 

to assessing the transient characteristics of the HRR calorimeter without the complicating 

effects of the fire spread and burning of real objects. One of the key results is the 

determination of the minimum pulse width needed to make an accurate measurement of 
34

peak HRR. For fires with FWHH of 15 s or longer, the calorimeter overestimates the 

actual peak heat release rate value by 3 % on average. However, if the FWHH measured 

by the calorimeter is less than 11 s, the measured peak heat release rate value 

underestimates the actual peak heat release rate by 15 % or more. 

The burner was also key to assessing the repeatability performance of the calorimeter. 

The coefficient of variation, CV, for repeat measurements of the peak heat release rate 

(instantaneous value) of fire pulses with a duration of 15 s or greater is typically less than 

0.02 for nominal 0.5 MW, 1.0 MW, and 2.5 MW fires. Such information is important in 

fire tests to separate calorimeter effects from the burning characteristics of real objects. 

It was found that the calorimeter is able to provide a good estimate of the total heat 

released for square-wave pulses changing too rapidly for the peak to be resolved by the 

calorimeter. That is, the integrated heat release measured by the calorimeter conserves 

energy to within about 5 % for pulses as short as 5 s. 

The fast response time of Schmidt-Boelter type radiometer was key to obtaining the most 

quantitative data on the gas burner output. It may be possible to combine the output of a 

radiometer, which has a time response of less than 1 s, with the integrated output from the 

calorimeter, which conserves energy, to estimate peak heat release rates for fires with 

FWHH as small as 3 s. 

2.4.3 Factors Affecting Dispersion of the Gas Measurement Signals 

As noted previously, the rise time of the O2, CO and CO2 signals depends not only on the 

inherent instrument response time14 but also on the time-wise dispersion of the gas 

concentration gradient due to mixing in the plume capture and gas sampling systems. 

Thus, the overall system response time assessed in the previous section is a product of 

various dispersion processes as well as instrument response characteristics. 

A transient fire under the calorimeter hood results in some time-varying stream of gas 

concentrations emerging from the flame tips that carries information about fire size (heat 

release rate). As this stream of gases enters the hood, it is paralleled by (and partially 

mixed with) a flow of in-drawn excess air. A series of lateral, cross-sectional slices 

through these parallel streams, with the gases mixed laterally15, contains a perfect history 

of the heat release rate behavior of the fire below the hood. However, no real system is 

capable of mixing these incoming gas streams in a lateral direction only. Invariably there 

is longitudinal mixing as well and this means that the time history of the fire is being 

smeared out. 

 

14 Instrument response time may be limited by the electronics and/or by the gas turnover time in the 

instrument’s plumbing and measurement cell. The latter can be shortened by an increased sample flow rate 

but this is typically limited by pressure drop or other mechanical issues. 

15 Lateral mixing (were it alone possible) brings with it only a lowering of the system sensitivity to heat 

release rate since it dilutes the fire plume gases. 
35

There are numerous longitudinal mixing mechanisms in any real system; each causes the 

process referred to as dispersion. The mechanisms pertinent to the six-meter hood system 

can be listed as follows: 

• Fire plume and duct flow turbulence 

• Smoke build-up in the hood 

• Flow separation in the bends of the hood duct 

• Flow dispersion in the sampling line 

• Mixing in the sample line filter and pumps 

• Dilution from side arms on a sampling line 

• Particle-induced mixing in the dry ice water trap and in the Drierite trap 

If the fire is enclosed in a room below the hood, a configuration that is sometimes used 

with the six-meter hood, there will be an additional significant mechanism due to plume 

mixing in the room upper hot gas layer prior to emergence from the room. Assuming that 

the upper layer in such a room is well-mixed, this room-effect dispersion time is of the 

order of the upper layer volume divided by the plume volumetric flow rate. The upper 

layer volume is at least as great as the volume determined by the door soffit height and 

will be greater with increased heat release rate. A program such as CFAST [3] can 

provide a good estimate of the depth of the hot gas layer. CFAST agrees well with 

experimental results for the ISO 9705 room (2.4m by 3m by 2.4 m high with normal size 

door opening) [19]. The plume flow rate depends on heat release rate as well; Heskestad 

[11] provides the equations for this calculation. As an example, a 500 kW fire in an ISO 

9705 room will have an estimated upper layer mixing time of 4 s to 8 s.16 This indicates 

that room dispersion effects can be substantial; as will be seen below, there is only one 

other element in the 3 MW system that normally yields comparable dispersion.17

We briefly consider the first item in the above list, which includes two dispersion 

mechanisms over which the experimenter has no control. The dispersion in the fire 

plume itself is a result of the intense turbulence typically existing there. For a 1 MW fire 

under the six-meter hood, one can show that the velocity-induced dispersion is of the 

order of ± 0.3 s in going from the top of the flames to the hood inlet. As to the second 

aspect, for a hood flow of 14.2 m3

/s, typical of that used for a large fire, the dispersion in 

the turbulent duct flow is of the order of ± 1 s (this dispersion mechanism is discussed 

further below in connection with the sampling line). 

Here we elaborate further only on those mechanisms over which an experiment designer 

might want to try to exercise some control. 

Confining the smoke build-up in the hood (including the skirted volume) might be 

thought of only as a safe way in which to handle a fire that temporarily becomes larger 

 

16 An object yielding a 500 kW fire, when placed in a room will also see substantial radiative feedback 

from the hot smoke layer. This could significantly increase the peak HRR seen in a room versus that in an 

open hood, depending on the view factor the burning surfaces have toward the upper smoke layer. 

17 One implication of this magnitude of room effect is that it has the potential to somewhat slow the overall 

system time response characteristics discussed in the above section on transient effects on HRR. 
36

than anticipated. In fact the skirt serves precisely this purpose but it also has 

consequences with regard to dispersion of which one should be aware. The skirt acts 

only by allowing a volume in which smoke can accumulate. Once the accumulation 

occurs to some depth, the entrainment of excess air is decreased only because the plume 

begins to entrain and re-cycle the air/smoke mixture in the backed-up volume, rather than 

further excess air (above the smoke layer base height). Thus the mechanism by which a 

hood skirt works is inherently dispersive. However, simulation of this process using FDS 

shows that the accumulated smoke volume is not well-stirred with the incoming plume. 

Instead the bulk of a fire plume (if centered on the hood duct inlet axis) will pass straight 

through the back-up volume, unchanged. Only the outer portions of the fire plume are 

sheared off and subjected to this re-cycling. A consequence of this is that the dispersion 

process is weak, mixing only a minority of the flow, but extended in time (many 

seconds); therefore, the accumulated smoke volume is returned to the hood duct flow at a 

slow pace. The system user should try to avoid having the plume partially back up into 

the hood/skirt volume if precise monitoring of peak heat release rate is desired. This 

requires having a sufficient hood flow to swallow the diluted fire plume at its peak (see 

Ref. [10] for a discussion of hood flow calculations based on plume entrainment). 

Flow dispersion in tube flow (both laminar and turbulent) is the subject of classic studies 

by G. I. Taylor [20]. Taylor shows that an instantaneous plug of fluid with a constant 

concentration of a gas species introduced into a fully developed turbulent flow results 

(after about 100 diameters of tube length) in a Gaussian distribution of that species, in 

time (or equivalent space). The width of the distribution is characterized by a dispersion 

coefficient that is a function of tube diameter and Reynolds number. Interestingly, 

Taylor’s solutions indicate that even in the ca. 25 m sampling lines between the roof and 

the instruments in the control room, turbulent flow dispersion has a negligible effect (less 

than 0.1 s) on system dispersion, even when the added effects of flow bends are 

accounted for. The same reference was used for the estimate above of the more 

significant dispersion in the hood duct flow. 

Dispersive mixing in such system elements as the sample filter upstream of the roof 

pump or the filters upstream of the control room pump(s) depends on the filter housing 

design. Such mixing would be eliminated if there are no separation zones for the flow 

but this tends to be incompatible with minimization of the size of the housing. 

Estimation of the possible mixing time constants (estimated flow separation volume over 

volumetric flow rate) for the existing filter housings implies that they are negligible 

sources of dispersion (less than 0.1 s) in the current system. The same applies to the 

sample pumps. 

Any location in the sampling system that has a pipe tee to provide an alternate flow path 

(or a tee to facilitate a pressure gage)18 introduces a side arm with a dead volume. 

Inspection of Figure 6 indicates that there are such tees in the present system. Such a 

dead volume communicates with the main sample flow by a combination of diffusion and 

 

18 Pressure gages do not need a significant flow in/out of their lead tube. Thus the lead tubes used in this 

system were made extremely small in diameter (1.5 mm OD tubing) in order to minimize their flow 

dispersive effects on the passing sample flow. 
37

induced flow circulation. Thus the gas in such a volume mixes with the passing sample 

flow, altering its composition and acting, in effect, as a dispersive mechanism. Using the 

pure diffusion solution for a semi-infinite, stationary gas subjected at time zero to a 

differing gas concentration at its outer boundary [21] one can show that pure diffusion 

can only cause mass fluxes that would amount to a negligible source of dispersion. On 

the other hand, the flow past the side arm can be expected to induce a circulation eddy in 

the outer end of the side arm of length comparable to the side arm tube diameter. This 

outer induced eddy will in turn induce another eddy of comparable dimensions. 

Presumably this induction continues into the side arm to a depth dictated by viscous 

dissipation of these induced eddy motions. Successive eddies communicate by diffusion 

and some flow mixing thus extracting gas from the side arm at a potentially much greater 

rate than by pure static diffusion alone. While there is not a quantitative treatment of this 

process available, one can anticipate that it could become a significant source of 

dispersion, especially for larger diameter side arms (worst case example, the side arm on 

the hood flow duct seen in Figure 2). Such dispersion could manifest itself as a drift in 

the gas analyzer outputs or it could extend the time to zero the gas analyzers. There is 

one location in the gas sampling system (in the control room) where this becomes a 

concern and the operating procedure therefore calls for closing a valve near the tee in this 

sidearm to greatly shorten its effective length and to minimize this effect. This action 

should eliminate this source of dispersion as an area of concern in the present system. 

The dry ice trap and the Drierite traps are both examples of packed beds. Flow 

dispersion in packed beds is a well-studied subject [22]. The complex flow of gas around 

the particles in a packed bed creates numerous paths of varying lengths along with minute 

separation zones, all of which spread the gas out in space (or time). The effect on a step 

function change in flow concentration as a result of axial dispersive processes in a 

successive set of particle beds like those used here for drying the sample flow can be 

reasonably well estimated. Given the particle size, gas flow rate and the length of the two 

drying devices, one can estimate that the combination smears an input step function in 

gas concentration over a time of about 4 seconds. These two drying devices are thus the 

most significant and routine sources of flow dispersion in the current system, though they 

do not add excessively to the basic system time response as set by the gas analyzer 

instruments themselves. If it became desirable to lessen the system response time further, 

it would be necessary to abandon the use of packed beds for gas flow drying. The Cone 

Calorimeter, for example, uses a laminar flow, thermoelectric cooling device. The 

sample flow is much greater here; it is not known whether an adequate device of this type 

is available. 

The dominant response time in the gas sample analysis system is that due to the analyzers 

themselves. The oxygen analyzer, for example, has a nominal response time of 8 s (to 90 

% of input change). The above discussion indicates that, in the absence of a room as a 

container for the fire being measured and also in the absence of significant smoke backup in the hood, the dispersive elements in the system will add roughly 5 s to this, giving 

an overall time to response to a step change at the fire of about 12 s to 13 s (depending on 

the exact definition of response time one uses19 and ignoring flow lag times)20. This is in 

 

19 Most of the dispersive times are (1/e) estimates. 
38

reasonable agreement with the results described in the preceding section on transient 

behavior of the heat release rate measurement system. Again it should be noted that use 

of a room or the existence of smoke back-up in the hood will increase the response time 

significantly, though the former will be more noticeable than the latter. 

2.4.4 Flame Height vs. Heat Release Rate 

The discussion by Heskestad [11] provides the basic equations governing the flame 

height from a pool fire, as a function of pool diameter. (Also included in the reference are 

other correlations for some other fire situations.) Since it is undesirable to allow the 

plume to be swallowed by the duct at the top of the hood, these flame height calculations 

provide one guideline for the system design, namely how far above the floor (the 

assumed fire location) the inlet duct needs to be in order to just capture the biggest fire 

for which the system is designed. The routine user of the system can also use such 

calculations to estimate whether the hood height will accommodate the planned fire. 

Since all fires do not act like simple pool fires21, one must use these calculations as a 

rough guide only and not push the upper limit without being prepared for some plume 

swallowing22. 

Flame heights have been calculated to assist in the design of experiments and 

measurements using the various exhaust hoods in the Large Fire Research Facility. The 

primary consideration here is assuring that the visible flame tip does not enter the exhaust 

system. Prediction of flame height for the calibration gas burner is difficult. At low flow 

rates, the flame emanating from the burner can be classified as buoyant, using 

Delichatsios’ Fire Froude Number [23]. However, with increasing gas flow the flame 

quickly transitions to a momentum dominated jet. Further complication of the problem 

results from flame merging. At low gas flow rates individual flamelets exist and then 

suddenly merge with increasing gas flow. It is unclear if the merging of individual 

flamelets occurs at the transition from a buoyant flame to a momentum driven jet. 

To predict burner flame height, it was assumed that the burner flames behave in 

aggregate like a pool fire. The results of the pool fire calculation are also instructive in 

judging likely flame heights from other burning objects, as well as from the calibration 

gas burner. Figure 17 shows the calculated flame tip height as a function of heat release 

rate for pool fires with diameters of 0.5 m, 1.0 m, and 2 m. Note that heat release rate is 

the primary determinant of flame tip height; pool fire diameter is secondary. The flame 

tip height is defined as the sum of the mean flame height and one-half of the 

 20 Also note that since all of these dispersive elements are in series, they are not strictly additive; rather the 

final result is the successive convolution of all preceding dispersive elements. 

21 Anything that impedes plume entrainment, such as a wall on one side of a fire, tends to increase flame 

height. Anything that tends to enhance plume entrainment, such as a multi-segmented fire will shorten the 

flame height. 

22 A small amount of plume swallowing will not likely suppress burning because the plume at its tip is 

over-aerated. This can be difficult to judge visually, however; a CO increase in the exhaust gases would be 

a better indicator. Since the gas temperature in the duct will go up rapidly if combustion continues there, 

there is a risk of damage to the system. Heavy swallowing of flames, especially for longer than times of 

the order of 10 s, is to be avoided. 
39

intermittency flame length. The intermittency length gives the difference between the 

maximum and minimum visible flame height, whereas the mean flame height is defined 

as the mean value of the visible flame height. The mean flame height and the 

intermittency lengths were taken from Eq. (8) and Figure 14 in Heskestad [24]. 

Calculation of flame height for multiple jets that merge, using Heskestad’s pool fire 

correlation, will over predict flame height. 

0.0

2.0

4.0

6.0

8.0

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

D=0.5m

D=1.0m

D=2.0m

Exp (NG Burner, Deff=1.32m)

Flame Tip Height, m

Heat Release Rate, MW 

6 m x 6 m Hood

Height from Floor

6 m x 6 m Hood

Height from Burner Surface

Figure 17 Calculated flame tip heights for pool fires and actual results for the 

natural gas calibration burner. 

The average flame tip height for a 6 MW fire in a 0.5 m burner is approximately 1 m 

higher than the average flame length. The plot shows two horizontal lines that represent 

the distance above the floor (solid line) and the distance above the calibration burner to 

the exhaust in the 6 m x 6 m (3 MW capacity) hood. Figure 17 shows that the 6 m x 6 m 

hood may not accommodate a 4 MW fire unless it has a diameter of 2 m or greater. Fires 

of 4 MW or greater and with effective diameters of less than 2 m are better 

accommodated by the 9 m x 12 m hood. 

Actual flame tip heights were measured when all eleven tubes were open. Flame tip 

height was determined using a 20 cm diameter protractor (1 degree angular resolution) to 

measure the angle between the line of sight to the flame tip and a horizontal reference 

line. The vertical position of the protractor above the floor and the horizontal location 

from the center of the burner was recorded. The height of the flame tip above the 

horizontal reference line is the tangent of the angle multiplied by the distance from the 

burner. Flame tip height above the floor (includes burner height) was computed by 
40

adding the vertical position of the protractor to the previously determined height. The 

measured flame tip heights minus the burner height are shown in Figure 17. The scatter 

is due to a combination of the variability in the use of the simple tools and the difficulty 

in estimating the point defining the flame tip, which fluctuates strongly. 

3 Operation and Safety 

3.1 Calorimeter 

The first steps to operating the calorimeter are to assure that the system is properly 

configured. This includes setting the exhaust hood flow to fit the anticipated fire size, 

bringing all components up to their normal operating temperatures, delivering the 

required flow from the sampling ports in the duct on the roof and feeding the correct 

flows to each of the analyzers. If any part of the gas analysis system has been opened 

(typically to insert clean water traps or particulate filters) it is checked for leaks. When 

everything is satisfactory, the data acquisition system (see Sec. 3.2 below) is initiated and 

calibration of the gas analyzers and zeroing of the pressure transducers on the bidirectional probes are performed. All analyzers are zeroed first using bottled nitrogen. 

They are then spanned with bottled calibration gases. After the system has fully 

recovered from the calibration process and is reading steadily once again the gas sample 

from the duct (with no fire), the no-fire baseline reading of the system is recorded by the 

data acquisition system. When the actual test is initiated, the baseline is again followed 

for three to five minutes before the fire is ignited (ignition is denoted with a marker 

channel). After the fire has extinguished naturally (or with an extinguishing agent), the 

post-test, no-fire reading of the system is recorded for at least three to five minutes. 

These before- and after-fire readings give information used to construct the system 

baseline that applies during the test. If the fire is located in an enclosure, more time is 

allowed for smoke to clear and for the system to equilibrate. 

The preceding steps involve various manipulations of the valves and switches on the 

control panels. Detailed and stepwise instructions are described in NISTIR 7052 [25]. 

3.2 Data Acquisition System 

The first series of steps to running the data acquisition system is to make sure that the 

computer and DAQ hardware have power and that the Measurement and Automation 

Explorer and LabView software packages are running. The DAQ program is then opened 

with LabView. Once the DAQ program is loaded, adjustments to operating parameters 

are made (these changes are made with the program loaded, but not running). The 

channels to be scanned are set23, the appropriate set of delay times for the hood flow is 

selected24, span values (mainly gas cylinder volume fractions) are set, and calorimetry 

settings for the combustion products expansion factor and heat of combustion per mass of 

 

23 Generally, the calorimetry channels are well established and changes to the channel list would involve 

non-calorimetry-related measurements. 

24 If a hood flow other than the three offered is required, the channel delay times would have to be changed 

(and possibly measured first). 
41

oxygen are entered. Particular chart history settings (the number of points retained on a 

chart) and the scan rate or averaging period (the default is 200/s and 1 s, respectively) 

must be entered before running the program. Ambient humidity and barometric pressure 

should be entered at this time and user-selected channels for display may also be set. 

These can be changed during the test as well. 

At this point, the program is started, the main filename for the test is entered, and the 

calorimetry system is operated as described in Sec. 3.1. After the fire is ignited, the DAQ 

operator monitors conditions and performance and updates the humidity and barometric 

pressure periodically over the duration of the fire test. After the fire is extinguished, the 

system is allowed to return to ambient conditions to record a post-fire baseline. The 

DAQ operator also uses the event marker channels to mark events such as zeros, spans, 

ambient conditions, baseline HRR, ignition, extinguishing, and any other significant 

occurrences. 

A detailed and stepwise description of the DAQ operating procedures is described in 

NISTIR 7052 [25]. 

3.3 Calibration Burner 

It is recommended to operate the calibration burner prior to a series of fire tests for 

operational confirmation. Operation of the burner requires a minimum of two operators 

at all times, the safety officer and the DAQ operator. The safety officer is responsible for 

all manual tasks involved in start-up and shut-down of the burner, while the DAQ 

operator is responsible for all remotely controlled tasks such as specifying the burner 

output. 

First the manual shut-off valve for the natural gas is opened and checks are made for 

proper operation of the igniters, the photoelectric sensor, and the solenoid valves. Once 

proper operation is confirmed, flow is initiated through the pilot tubes and the pilots are 

ignited using the DAQ burner control panel. Upon successful ignition a target heat 

release rate is specified at the DAQ burner control panel. The DAQ program provides a 

current to the natural gas control valve which has been calibrated to corresponding heat 

release rates, therefore giving remote control of the output of the burner. The burner and 

the DAQ burner control panel are monitored as the test proceeds, with all personnel being 

instructed to warn for unsafe conditions. The burner is shutdown by reducing the target 

heat release rate so that flow occurs through the pilot tubes and shutting off the flow 

control valve. The pilot flames eventually extinguish and the natural gas line is purged 

from the shut-off valve with nitrogen. 

Detailed operation steps are described in NISTIR 7502. It is important to follow these 

procedures to ensure safe operation of the calibration burner. Improper operation of the 

burner could produce an explosive mixture of natural gas and air within the building, 

possibly resulting in a catastrophic event. Emergency shut-down procedures are also 

described in NISTIR 7502. 
42

3.4 Safety 

A fire of any size, whether controlled or un-controlled, is inherently hazardous. 

Therefore, when dealing with large fires, safety is a priority concern. Many active and 

passive safety mechanisms are in place in the Large Fire Research Laboratory. The most 

important of these mechanisms is allowing only adequately trained and experienced 

personnel to operate and participate in fire tests. Safety steps have been documented as 

part of the operations procedures for many of the components and tasks involved in a fire 

test. A fire test requires at least three personnel to comprise a test team. Each member of 

the testing team has a specific role and is required to update an operations checklist prior 

to beginning a fire test. The safety officer conducts a short meeting of the team members 

prior to the fire test to explain the test and reconfirm each member’s responsibility. 

During this meeting each operations checklist is collected and reviewed by the safety 

officer to ensure that no safety steps are omitted. 

Other safety mechanisms include keeping the fire test area cleared for visual monitoring, 

employing communication devices to convey warning signals and event changes, having 

the properly sized water extinguishing system ready and, if necessary, having a trained 

and fully equipped firefighter on hand. There are safety systems to manage unforeseen 

failures of components like the natural gas burner such as manual shut-off valves and 

emergency shut-off buttons located near exits. Also the facility exhaust system is 

equipped with two exhaust trains to reduce the risk of complete failure of the system 

during a fire experiment. Finally, routine maintenance of the facility and its components 

is performed and safety procedures are reviewed annually. 

3.5 Baseline Heat Release Rate 

There is virtually always some drift in the gas analyzers, especially the oxygen analyzer. 

As a result there will be a non-zero (positive or negative) baseline measurement of heat 

release rate. This indicated value of heat release rate will exist before the fire test has 

started and after the fire (including smoldering) is fully extinguished (and the pre and 

post-test values will usually differ). The technique used to account for the baseline drift 

begins with allowing the heat release rate facility to operate for (3 to 5) min immediately 

prior to a fire test and (3 to 5) min immediately following full extinguishment of a fire 

test. The pre-test and post-test baselines are therefore determined, respectively. Figure 

18 is a graphical example of a fire test performed with the calibration burner. Nominal 

heat release rates were 0.05 MW, 0.65 MW and 2.70 MW and a baseline drift from 0.005 

to 0.020 MW was observed over the span of 54 minutes. 
43

0.0

1.0

2.0

3.0

0

0.01

0.02

0.03

0.04

2550 3400 4250 5100

HRR

Background

HRR, MW

Background HRR, MW

Time, s

Figure 18 Time history of uncorrected heat release (left axis) and baseline heat 

release rate (right axis). 

A time-varying relation for the baseline heat release rate, qbkgd ,sys & , relative to which all 

heat release rate measurements during a test were made, is determined by performing a 

least squares linear fit to the pre- and post-test data, Eq. (14) (see Figure 18). This 

relation assumes that the baseline drift of the facility is linear over the entire duration of 

the fire test. (30 minute drift test of the gas analyzers resulted in mostly linear positive or 

negative drift, therefore confirming this assumption) The estimated baseline heat release 

rate at time t is algebraically subtracted from the measured value of heat release rate at 

time t, Eq. (15). Performing this operation at each time step results in a time history of 

heat release rate corrected for the baseline drift of the system, qcrctd ,sys & . This operation is 

especially important when evaluating peak heat release rate. The baseline will affect the 

characteristics, such as height and width, of the peak heat release rate curve. 

bkgd sys sys sys q = m t + j , & ( 14 ) 

qcrctd ,sys q qbkgd ,sys & = & − & ( 15 ) 

The method outlined above measures the drift of the system as a whole. It will be 

referred to as Method I. The system drift is a result of the drift of the system 

components. 

Since the raw data from each instrument is recorded for the pre- and post- test 

measurements, the baseline heat release rate due to the component drift also can be 

computed. The heat release rate is derived from the measured depletion of oxygen, 
44

which is described mathematically (Eq. (7)) by the oxygen depletion factor, φ . 

Measurements of CO2 and CO are also used to define φ . For the present report, the 

baseline drift of the three gas analyzers (O2, CO2, and CO) is used to compute a baseline 

drift of heat release rate from conditions at time, t=0. Least squares linear fits of the 

voltage outputs of each instrument, Eq. (16) to (18), are used to compute the gas species 

volume fraction drift. The baseline volume fractions are substituted into Eq. (13) to 

compute the baseline heat release rate due to the instrument drift. 

Computing the baseline heat release rate due to the drift of the components will be 

referred to as Method II, and is represented by Eq. (19). Baseline heat release rates are 

computed by both methods for two series of test at random times within a series. The 

results are presented in Table 4 and show good agreement for the two methods of 

baseline HRR determination. The fractional difference, Fd, defined as (Method I -

Method II)/Method I has a mean and standard deviation of 0.12 ± 0.10. This good 

agreement indicates that most of the background drift can be accounted for by the drift of 

the gas analyzers. 

2 2 2 O ,bkgd VO VO V = m t + j ( 16 ) 

2 2 2 , CO bkgd VCO VCO V = m t + j ( 17 ) 

CO bkgd VCO VCO V = m t + j , ( 18 ) 

 ,.....) (....., , , , , , , qbkgd comp q VO2 bkgd VCO2 bkgd VCO bkgd & = & ( 19 ) 

Table 4 Baseline heat release rate comparison 

Baseline heat release rate, kW 

System 

(Method I) 

Component 

(Method II) 

Fd 

14.8 11.6 0.22 

16.6 12.8 0.23 

17.4 11.6 0.26 

34.8 35.7 -0.03 

39.5 38.8 0.02 

51.9 50.2 0.03 

55.8 52.1 0.07 

65.1 57.1 0.12 

 Avg ± σ 0.12 ± 0.10 
45

4 Assessment of Uncertainty 

A few studies exist which address the uncertainty of heat release rate measurements by 

oxygen consumption calorimetry. The studies vary in apparatus, version of the heat 

release rate equation, level of heat release rate, and detail of analysis. Axelsson et al. [26] 

estimated the heat release rate measurement uncertainty for the measurement as 

conducted in the Single Burning Item (prEN 13823) and the Room Corner Test (ISO 

9705). The relative expanded uncertainty estimates ranged from 0.07 to 0.14, depending 

on the apparatus and the heat release rate. The uncertainty of the oxygen concentration 

measurement, followed by the heat of combustion factor and the mass flow rate 

measurement were identified as the major sources of uncertainty. Details on the 

uncertainty of the oxygen and the mass flow rate measurements were presented. The 

study notes that for larger oxygen deficits the combined uncertainty of the heat release 

rate measurement is less. Enright and Fleischmann [27] performed an analytical estimate 

of the heat release rate measurement uncertainty for the cone calorimeter. They 

estimated relative expanded uncertainty ranging from 0.10 to 0.12, depending on the heat 

release rate. The greatest sources of uncertainty identified were, the heat of combustion 

factor, the combustion expansion factor, and the oxygen measurement. Yeager [28] also 

performed an analytical uncertainty estimate for experiments conducted in a room 

compartment with a controlled supply of energy into the room and identified the volume 

flow rate and oxygen measurement as majors source of uncertainty. He reports a relative 

expanded uncertainty of 0.12 when the exhaust volume flow rate is optimized for the fire. 

The oxygen depletion measurement and the exhaust mass flow rate measurement have 

been consistently reported as major sources of uncertainty. However, Janssens’ [29] 

review of the results of heat release rate round robins brings attention to the uncertainty 

due to random effects such as material and burning variability, environmental effects, 

operator error, and measurement bias between laboratories. The impact of these effects is 

important and deserves further study. A crucial first step in such a study is a self-audit by 

the individual laboratories of their heat release measurement uncertainty. Such an 

assessment requires the use of a calibration burner with good repeatability to avoid the 

variability in the results from differences in the heat release rate for nominally identical 

solid fuel structures. The goal is to evaluate the calorimeter and one wants to avoid other 

sources of variability. 

The self audit conducted here follows the methodology described in the ISO guidelines 

[30] and adopted by NIST [31]. The approach is to represent each component of 

uncertainty as a standard uncertainty. So called Type A uncertainties are those that can 

be computed based on statistics such as the standard deviation about the mean. The 

others, designated as Type B, require scientific judgment together with available data. A 

common example is knowing the maximum range of a variable and then converting this 

range into a standard uncertainty by dividing it by 2√3 [31]. 

The calculation of the uncertainty in the heat release rate requires the combining of many 

components of uncertainty. This is apparent from the heat release rate equation, Eq. (13), 

which is a function of more than ten variables. There are five new aspects to the 

uncertainty analysis presented here as compared to previous analyses. The first concerns 
46

the estimation of the uncertainty in the oxygen volume fraction to take into account the 

small noise relative to a much larger uncertainty in the calibration gas. A second is 

expressing all of the gas analyzer outputs in terms of their voltage to eliminate 

correlations induced in intermediate quantities by common measurement results. A third 

is subtracting a baseline drift from the gas analyzers and analyzing the uncertainty in the 

drift correction. The fourth is to perform the uncertainty propagation with a numerical 

method that is easily implemented in a spreadsheet program. And the fifth is to combine 

the total uncertainty based on the uncertainty propagation with the repeatability 

measurements to obtain a combined uncertainty and ultimately an expanded uncertainty 

corresponding to an approximate 95 % confidence interval. The assessment begins with 

estimating the uncertainty of individual Type A and B components. 

4.1 Type A Uncertainty 

Voltage

The standard uncertainty is estimated using the sample standard deviation of data 

collected during periods of operation at constant heat release rate or without any heat 

input into the system. This applies to all gas analyzer voltages and pressure transducer 

voltages. For each analyzer there are at least three voltage measurements. First, the 

voltage corresponding to the zero gas (nitrogen), second the voltage corresponding to the 

span gas, and last the voltage corresponding to the gas volume fraction at time t. For 

example, the voltage for the oxygen analyzer corresponding to a volume fraction of 

0.2095 was approximately 0.833 V and the uncertainty in the voltage reading by the 

DAQ was ±25 μV. 

Temperature

Type K thermocouples are employed for each of the temperature measurements. These 

thermocouples have fundamental error limits of ±2.2 K. Assuming a rectangular 

distribution, the standard uncertainty is computed by dividing the limit by √3. The 

thermocouples are directly in the exhaust stream and are subject to fluctuations due to the 

turbulence in the exhaust duct. The standard deviation of 300 temperature measurements 

is nominally 1.0 K and is chosen to represent the noise due to turbulence. Combining the 

two standard deviations in quadrature results in a standard uncertainty of ±1.6 K. 

Repeatability

The mean and the standard deviation about the mean were computed for the six repeat 

heat release rate measurements for the nominal 0.05 MW, 0.65 MW, and the 2.70 MW 

natural gas burner flames. 

Linear Least Squares Coefficients

The system and instrument drift data typically consist of 300 to 800 points. An Excel 

spreadsheet function “LINEST” is used to compute the coefficients for the linear fit. It 

also computes the standard deviation of the mean (or standard error) for each coefficient 

which is referred to here as the standard uncertainty. Before the regression analysis is 

run, the data are centered about the mean time, t , which eliminates the covariance 

between the slope and the intercept. 
47

4.2 Type B Uncertainty 

Zero and Span Gases

The zero gas is 99.99 % nitrogen, therefore the measurements of zero O2, CO2, and CO 

volume fractions are assumed to have a standard uncertainty of ±0.0001. The O2 span 

gas is breathing air that has an oxygen volume fraction of 0.2095 ±0.0005. Span gases 

for CO2 and CO are purchased as a mixture of CO2/CO/N2. The volume fractions are 

specified and each bottle is delivered with a certificate of analysis specifying the actual 

volume fraction with a relative standard uncertainty of ±0.005. 

Relative Humidity

Based on the manufacturer’s technical specifications the accuracy of the humidity meter 

is ±2 % RH. 

Ambient Pressure

The ambient pressure is read from a mercury barometer and corrected for temperature 

and altitude. The uncertainty in the reading is estimated to be ±0.27 kPa. 

Bi-Directional Probe Constant, Cbdp,i 

For flow with Reynolds number (relative to probe diameter) greater than 3800, the ratio 

of dynamic pressure to velocity is essentially constant and this constant value (0.0698 

m3

/K-kg) has been determined to be accurate to within ±5 % (±0.0035 m3/K-kg)[13]. 

The nominal range of Reynolds number for flows over the 2.54 cm bi-directional probes 

is 2000 to 16000. 

Duct Diameter

The duct diameter, 1.52 m, is measured with a tape measure. Due to the duct not being a 

perfect circle the standard uncertainty is assumed to be ±0.01 m. 

 

Heat Release of Fuel per kg of Oxygen Consumed, HC Hc Mass _ O2 (∆ )

The value of HC Hc Mass O2 _ (∆ ) for use if the fuel properties are not known is taken to be 13.1 

MJ/kg O2 based on the study of Huggett [5]. A slightly modified version of Table 2 of 

Huggett for typical synthetic polymers is used to compute the standard deviation in 

HC Hc Mass _ O2 (∆ ) , ±0.35 MJ/kg O2. The modifications were to delete polyoxymethylene, 

which is an outlier with an HC Hc Mass _ O2 (∆ ) of 14.5 MJ/kg O2, and to add cellulose with an 

HC Hc Mass O2 _ (∆ ) of 13.59 MJ/kg O2, which is representative of paper. The relative standard 

uncertainty is computed as ±0.027. 

Natural gas, which contains typically 94% methane (as supplied by the local natural gas 

company) is used for calibration burns. Fuel composition is periodically determined 

from gas chromatograph results. The value of HC Hc Mass _ O2 (∆ ) for natural gas is 12.55 

MJ/kg O2 with an estimated standard relative uncertainty of ±0.00159 (±0.02 MJ/kg O2). 

It is computed from the enthalpy balance for the complete combustion of natural gas 
48

(major species) and oxygen to produce CO2 and H2O. (The complete combustion of pure 

methane and oxygen, which is a simpler computation of the enthalpy balance, results in 

HC Hc Mass O2 _ (∆ ) = 12.54 MJ/kg O2.) The NIST Chemistry WebBook lists the heat of 

formation for each molecule and the expanded uncertainty [32]. 

Heat Release of CO per kg of Oxygen Consumed, CO Hc Mass _ O2 (∆ ) 

The value of CO Hc Mass O2 _ (∆ ) , computed from the enthalpy balance for the combustion of 

CO in oxygen to produce CO2, is 17.69 MJ/kg O2 with an estimated standard relative 

uncertainty of ±0.0006 (±0.01 MJ/kg O2). The NIST Chemistry WebBook lists the heat 

of formation for each molecule and the expanded uncertainty [32].

Combustion Products Expansion Parameter, α

The expansion parameter α accounts for the volume of the combustion products 

exceeding that of the oxygen consumed. The quantity α can be related to the ratio of the 

moles of combustion products formed to the moles of oxygen consumed, which is 

denoted as β, via the following equation: 

1 ( 1) 2 α = + β − o X O ( 20 ) 

where o X O2 is the volume fraction of oxygen in ambient air, which is assumed to be 

0.2095. As an example, the combustion of one mole of methane (CH4) requires two 

moles of oxygen and produces three moles of product made up of one mole of CO2 and 

two moles of water. The resulting values of β and α are 1.5 and 1.105. 

The values of α for a range of common polymers found in constructed facilities are listed 

in Table 5. The mean value of α, 1.10, is used in the computation of the heat release rate 

if the properties of the fuel are unknown. In this case, the standard uncertainty for α is 

computed as the standard deviation of the values for the polymers listed in Table 5 with a 

value of ±0.048 and a standard relative uncertainty of ±0.044. 

Table 5 Expansion parameter α for polymers found in constructed facilities 

Polymer Formula β α 

Polystyrene (-C8H8-) 1.200 1.042 

Polycarbonate (-C16H11O3-) 1.246 1.052 

Polyethylene (-C2H4-) 1.333 1.069 

Polymethylmethacrylate (-C5H8O2-) 1.500 1.105 

Polyvinylchloride (-C2H3Cl-) 1.600 1.126 

Polyurethane (-C13H6O6N2-) 1.630 1.132 

Cellulose (-C6H10O5-) 1.830 1.175 

Avg ± σ 1.100 ± 0.048 
49

4.3 Uncertainty Propagation 

Suppose a general equation can be applied to describe a measurement process which has 

an output y and the output y is based on a number of input quantities, xi. 

 ) ( , , ,...., 1 2 3 N y = y x x x x ( 21 ) 

In the case that all input quantities are uncorrelated, the combined relative expanded 

uncertainty is given by 

∑= 













 = =

N

i i

i

i

e c

x

u x k s

y

u y k

y

U y

1

2

2 ( ) ( ) ( ) ( 22 ) 

In Eq. (22), u(xi) is the standard uncertainty for each input, uc(y) is the combined 

uncertainty, k is the coverage factor, and si is the associated dimensionless sensitivity 

coefficient given by 

y

x

x

y s i

i

i ∂

∂ = ( 23 ) 

4.3.1 Calibration Burner 

An estimate of uncertainty for the natural gas burner heat output, ut( qburner & ), is obtained 

by applying the propagation of uncertainty as described in Eq. (22). This includes the 

uncertainty of the three measured variables; namely pressure (P), temperature (T) and 

volumetric flow rate (V& ) and the uncertainty of the natural gas heat of combustion 

( Hc NG (∆ ) ). The ideal heat output is computed from the following equation: 

c NG

ref

ref

burner H

T

T

P

P

q = V (∆ ) & & ( 24 ) 

Volumetric flow, temperature and pressure were measured and the volume flow rate was 

adjusted to represent the standard conditions for which the heat of combustion was 

determined. Since the ideal heat output is the multiplicative product of each 

measurement, the sensitivity coefficients reduce to 1.0 (or -1.0) and the resulting 

expression for the uncertainty is given in the following equation: 

2

( )

2 2 2

( )

( )

















∆  + 

 



 + − 













  + 

 



 = ∆

c NG

P V T H

burner

t burner

H

u

T

u

V

u

P

u

q

u q c NG

& &

& & ( 25 ) 

The Type K thermocouple used to measure the natural gas temperature has error limits of 

±1.1 K, which corresponds to a standard uncertainty of ±0.64 K when a rectangular 
50

distribution is assumed. The uncertainty of the OmegaDyne pressure transducer is 0.10 

% of the full range, which is 10.2 atm resulting in a standard uncertainty of ±0.0102 atm. 

Nominal temperature and pressure are 293 K and 2.02 atm, respectively. The lower 

heating value of the natural gas heat of combustion was determined from gas 

chromatograph reports supplied by the local gas utility. The nominal value is 33,733 

kJ/m3

 (computed for reference conditions of 1 atm and 300 K) and the standard deviation, 

computed from gas chromatograph reports spanning six days, was ±133 kJ/m3. Neither 

of these three variables are expected to fluctuate significantly during a test. The 

volumetric flow rate was the only dynamic variable in this calculation. It is determined 

from the ratio of the measured frequency of pulses, f, and the meter factor, K. The meter 

factor supplied by the manufacturer, K=5.08x10-4 pulses/cm3

, is employed in the 

computation. This agrees with the mean meter factor taken from the results of the flow 

calibration [25]. The standard deviation about the mean meter factor over the calibration 

range is ±0.03x10-4 pulses/cm3

. Combined with the uncertainty in the pulse counting 

method, the standard relative uncertainty of the volumetric flow rate measurement is 

estimated at ±0.01. The standard relative uncertainty for the temperature, pressure and 

natural gas heat of combustion are, ±0.002, ±0.005, and ±0.004, respectively. Using Eq. 

(25), the total relative uncertainty for the ideal heat output of the calibration burner is 

calculated to be ±0.012. 

4.3.2 Calorimetry 

The relationship which applies to the measurement of heat release rate is given by Eq. 

(13), which is reproduced here for convenience. 

o

O

air

O

o

e H O

O

HC CO

c Mass O

CO

c Mass O

HC

c Mass O

X

M

m X M

X

X

q H H H

2

2 2

2

2 2 2

(1 )

1 ( 1)

] 2

1 [( ) (( ) ( ) ) _ _ _















 −

+ −

− = ∆ − ∆ − ∆

φ α

φ φ

&

&

 (13) 

with 

o

O CO CO O

o

CO CO O CO

o

O

X X X X

X X X X X

2 2 2

2 2 2 2

(1 )

(1 ) (1 )

− − −

− − − − φ =

As written the expression for q& contains parameters which are correlated or codependent, namely O2 X and o X O2 , which violates the assumption leading to Eq. (22). In 

order to avoid such co-dependencies induced by common inputs, Eq. (13) must be 

reduced to the most basic measurement inputs of instrument voltages, thermocouple 

temperatures, and constant parameters (universal, empirical and calibration). This 

process is described for each of the measurement devices in text to follow. 

The resulting measurement value for the project at hand is the heat release rate corrected 

for the baseline or zero measurement bias, the algebraic subtraction of Eq. (13) and (19). 
51

The baseline heat release rate is determined by linear fits to the component data (Eq. (19), 

Method II). For the purpose of the uncertainty propagation, all data is centered about the 

mean time, t (determined from the baseline data), which eliminates the covariance 

between the slope, m, and the intercept, j. 

a = m(t − t) + j ( 26 ) 

It is important to recognize that the quantity φ is proportional to ( o X O2 - O2 X ) since the 

volume fraction of CO and CO2 are small compared to that of oxygen. In addition the 

difference in the oxygen volume fractions is typically quite small. Some care must be 

taken in estimating the uncertainty in this case. An example is presented by Kragten [33] 

which illustrates the potential danger caused by ignoring correlations induced in 

intermediate measurement results by shared inputs. A similar example affecting the 

assessment of heat release rate is appended. The example illustrates the importance of 

performing an uncertainty analysis for the most basic inputs such as instrument voltage 

outputs and constants (empirical or universal). The relations for converting the 

instrument voltage outputs to units of more scientific significance follow. 

Oxygen Analyzer

The oxygen analyzer produces a DC voltage [V] that is converted to the oxygen volume 

fraction by the following relation (includes the pressure correction): 















 − −

− =

O press

O press Zero

O Span O Zero

O Span O Zero

O O Zero

O V

V

X X

V V

V V

X

,

, ,

, ,

, ,

,

2

2

2 2

2 2

2 2

2 ( ) ( 27) 

where: 

O2 V = voltage measured by oxygen sensor at time, t 

VO2 ,Zero = voltage measured by oxygen sensor for the zero gas 

VO2 ,Span = voltage measured by oxygen sensor for the span gas 

X O2 ,Span = oxygen volume fraction of the span gas, bottled air (0.2095) 

X O2 ,Zero = oxygen volume fraction of the zero gas, nitrogen (N2) 

O2 X = oxygen volume fraction computed with respect to voltage measurement at time, t 

VO , press 2 = voltage measured by absolute manometer to monitor pressure in the oxygen 

analyzer at time, t 

VO2 , press,Zero = voltage measured by absolute manometer to monitor pressure in the oxygen 

analyzer during the zeroing of the instrument 

The oxygen sensor is calibrated before each fire test to determine VO2 ,Span and VO2 ,Zero . 

Absolute pressure is measured at the oxygen analyzer and the output is corrected for 

pressure fluctuations relative to the output when the instrument is zeroed. 
52

Similarly, the relation converting the voltage output to the ambient oxygen volume 

fraction is: 















 − −

− = o

O press

O press Zero

O Span O Zero

O Span O Zero

O Zero

o

o O

O V

V

X X

V V

V V

X

,

, ,

, ,

, ,

,

2

2

2 2

2 2

2 2

2 ( ) ( 28) 

The ambient oxygen measurement is conducted immediately after the spanning of the 

instruments. This value is then held as a constant in the computation. 

Carbon Dioxide Analyzer

The carbon dioxide analyzer produces a DC voltage [V] that is converted to the carbon 

dioxide volume fraction by the following relation: 

( ) , ,

, , , ,

, ,

2 2

2 2 2 2

2 2 2 2

2 CO Span CO Zero

CO Span CO Span CO Zero CO Zero

CO CO CO Zero CO Zero

CO X X

V RF V RF

V RF V RF

X − −

− = ( 29 ) 

where: 

CO2 V = voltage measured by CO2 analyzer at time, t 

VCO2 ,Zero = voltage measured by CO2 analyzer for the zero gas 

VCO2 ,Span = voltage measured by CO2 analyzer for the span gas 

X CO2 ,Span = CO2 volume fraction of the span gas, CO2/CO/N2 mixture 

X CO2 ,Zero = CO2 volume fraction of the zero gas, N2 

CO2 RF = range factor for the CO2 voltage measured at time,t 

RFCO2 ,Zero = range factor for the CO2 voltage measured when the instrument is zeroed 

RFCO2 ,Span = range factor for the CO2 voltage measured when the instrument is spanned 

The range factors are specific to the CO2 analyzer employed. Depending on the range 

selected by the analyzer, a voltage is sent to the appropriate range channel of the data 

acquisition system. This voltage (VRi) is entered into the following relation to determine 

the range factor. The range factors are written to the adjusted data file. 

RFCO2 = AVR1 + BVR2 + CVR3 + DVR4 ( 30 ) 

Similarly, the voltage output converted to the ambient carbon dioxide volume fraction is: 

( ) , ,

, , , ,

, ,

2 2

2 2 2 2

2 2 2 2

2 CO Span CO Zero

CO Span CO Span CO Zero CO Zero

CO Zero CO Zero

o

CO

o

o CO

CO X X

V RF V RF

V RF V RF

X − −

− = ( 31 ) 
53

Carbon Monoxide Analyzer

The carbon monoxide analyzer produces a DC voltage [V] that is converted to the carbon 

monoxide volume fraction by the following relation: 

( ) , ,

, , , ,

, ,

CO Span CO Zero

CO Span CO Span CO Zero CO Zero

CO CO CO Zero CO Zero

CO X X

V RF V RF

V RF V RF

X − −

− = ( 32 ) 

where: 

VCO = voltage measured by CO analyzer at time, t 

VCO,Zero = voltage measured by CO analyzer for the zero gas 

VCO,Span = voltage measured by CO analyzer for the span gas 

X CO,Span = CO volume fraction of the span gas, CO2/CO/N2 mixture 

X CO,Zero = CO volume fraction of the zero gas, N2 

RFCO = range factor for the CO voltage measured at time,t. 

RFCO,Zero = range factor for the CO voltage measured when the instrument is zeroed 

RFCO,Span = range factor for the CO voltage measured when the instrument is spanned 

The range factors will be specific to the CO analyzer employed. Depending on the range 

selected by the analyzer, a voltage is sent to the appropriate range channel of the data 

acquisition system. This voltage (VRi) is entered into the following relation to determine 

the range factor. The range factors are written to the adjusted data file. 

RFCO = AVR1 + BVR2 + CVR3 + DVR4 ( 33 ) 

Ambient Water Vapor

Relative humidity (RH, [%]) is read from a humidity meter located in the data acquisition 

control room. The ambient temperature (Tamb, [o

C]) is measured from a thermocouple 

located on the data acquisition rack. The signal condition modules convert the 

thermocouple millivolt signal to degrees C before it is used in the HRR computation. 

Ambient pressure (Pamb, [torr]) is read from a mercury barometer located in the high bay 

area of the Large Fire Laboratory. The reading is corrected for altitude and temperature. 

The three measurements are used in the following relation [34] to infer the volume 

fraction of water vapor at ambient conditions. 

0.043725 0.611)

(5.5259 10 1.7195 10 1.5678 10

101.3

760

100

7 4 5 3 3 2

2

+ +

= + + − − −

amb

amb amb amb

amb

o

H O

T

x T x T x T

P

RH X

( 34 ) 

Exhaust Duct Bi-Directional Probes and Thermocouples

The bi-directional probes produce a DC voltage (Vbdp,i) that is converted to a differential 

pressure by the following relation: 
54

 ) [ ] 133.322 ( ∆Pi Pa = bi Vbdp,i −Vbdp,i,Zero ( 35 ) 

where bi [torr / V] is the calibration coefficient of each probe and 133.322 [Pa / torr] is 

the conversion factor for torr to Pascals. The voltage measured when each probe is 

zeroed is Vbdp,i,Zero. 

At each probe location a temperature is measured, Tbdp,i. The signal condition modules 

convert the thermocouple millivolt signal to degrees C. The pressure differential and 

temperature measurements are employed in the following relation to infer the exhaust gas 

velocity at each probe location. 

( 273.15) Ubdp,i = Cbdp ∆Pi Tbdp,i + ( 36 ) 

where 

0.0698 1 2 = =

ref air

bdp P MW

R

C

C ( 37 ) 

when C=1.08 ±0.05 as determined by McCaffrey and Heskestad [13]. 

The average velocity and average temperature for the six probes are computed to 

represent the velocity and temperature in the exhaust duct at the gas sampling point. 

∑=

= − +

6

1

, , , , 133.322 ( )( 273.15) 6

0.0698

i

Uavg bi Vbdp i Vbdp i Zero Tbdp i ( 38 ) 

∑=

= +

6

1

, ( 273.15) 6

1

i

Tavg Tbdp i ( 39 ) 

The mass flow rate ([kg/s]) of the exhaust duct is then computed by the following 

relation: 

4 ( 273.15) 6

1

133.322 ( )( 273.15) 6

0.0698

2

6 ,

1

,

6

1

, , , , D T

T

b V V T

m air ref ref

i

bdp i

i

i bdp i bdp i Zero bdp i

e

π

ρ

∑

∑

=

=

+

− +

& = ( 40 ) 

where ρair,ref (1.1767 kg/m3) is the density of the reference gas (dry air), Tref (300 K) is 

the temperature of the reference gas, and D (1.52 m) is the inside diameter of the exhaust 

duct at the probe location. 

A close inspection of all of the relations presented above will reveal the co-dependent 

input parameters of Eq. (13). The co-dependent pairs are ( O2 X , o X O2 ), ( CO2 X , o X CO2 ), and 

(Ubdp,i,Tbdp,i). These parameters share variables or one is dependent upon the other as in 
55

the case of Ubdp,i and Tbdp,i. Therefore it is recommended to break down the computation 

to the basic input measurements before beginning an uncertainty propagation to properly 

account for the co-dependence in intermediate quantities. 

A major justification for the added complexity to the uncertainty propagation is specific 

to the heat release measurement as conducted in this laboratory and is illustrated by the 

appended example. The heat release measurement is based on oxygen consumption 

calorimetry and is therefore directly proportional to the difference between the 

measurements of oxygen and ambient oxygen volume fraction, ( o X O2 - O2 X ). The 

combined uncertainty for each term of the oxygen difference is dominated by the 

uncertainty of the oxygen volume fraction of the span gas. This uncertainty is on the 

order of 3 times greater than the oxygen difference measured in small fires. By rewriting 

Eq. (13) in terms of the basic inputs, the uncertainty of the span gas will be correctly 

accounted for and will not give an erroneously large uncertainty for the oxygen 

difference, as would be the case if Eq. (13) were used directly. 

Rewriting Eq. (13) in terms of the basic input measurements results in a final equation 

that consists of 64 input parameters. The parameters are listed in Table 6 and the 

equations, in terms of basic inputs, for each measurement instrument were presented 

above. 

A numerical approach to Eq. (22) as described by Kragten [33] has the advantage of 

being easily implemented in an Excel spreadsheet. The method numerically 

approximates all the partial derivatives and calculates the combined standard uncertainty 

without the use of intermediate quantities that violate the condition of mutual 

independence. It is a spreadsheet-based format that offers relatively easy tracking of the 

contribution of uncertainty of each parameter. An example of the spreadsheet developed 

for this project is shown in Figure 19. The details of creating a spreadsheet uncertainty 

propagation are explained in Ref. [33]. 

The uncertainty of the heat release rate measurement varies with the operating conditions 

of the calorimeter and the size of the fire. A series of experiments was conducted with 

the natural gas calibration burner targeted at 0.50 MW, 0.65 MW, and 2.70 MW, to 

produce nominal values for the uncertainty analysis. The measurements were corrected 

for the background drift by applying Method II as described in Sec. 3.5 and the 

uncertainty propagation was performed on the resulting relation for the corrected heat 

release rate. Due to the background correction the number of parameters increases from 

64 to 70. 
56

Figure 19 Example of Kragten spreadsheet for computing HRR uncertainty. 

The uncertainty for these measurements is presented in Table 6 including the total 

relative uncertainty, ut. Each of the 70 input parameters is listed with its nominal value, 

relative standard uncertainty and its non-dimensional sensitivity coefficient. The 

quantities labeled as “Constant” are quantities that may or may not vary between uses of 

the calorimeter while those labeled “Variable” fluctuate during use of the calorimeter. 

Type A or B determinations are indicated by either “A” or “B” in parenthesis. By visual 

inspection of the table, the product of the sensitivity coefficient and the relative standard 

uncertainty help to quantify the contribution to the total uncertainty. Therefore 

parameters with absolute values of sensitivity coefficients equal to or greater than unity 

are expected to make a greater contribution to the total uncertainty. 
57

Table 6 Corrected heat release rate relative uncertainty 

Sensitivity coefficient, |si| Parameter/measurement 

[units] (type) 

Nominal 

value 

Relative 

standard 

uncertainty, 

u(xi) / xi

0.05 

MW 

0.65 

 MW 

2.70 

MW 

O2 X , o X O2 , [-] 

O2 V , [V] (variable) 0.817985 0.00003 (A) 1113 80.2 16.35 

VO2 ,Zero , [V] (constant) -0.002900 0.00869 (A) 0.004 0.004 0.004 

VO2 ,Span , [V] (constant) 0.832965 0.00003 (A) 1.253 1.256 1.24 

X O2 ,Zero , [-] (constant) 0.0000 0.0001 (B) 

absolute 

- - - 

X O2 ,Span , [-] (constant) 0.2095 0.00239 (B) 1.258 1.261 1.245 

VO , press 2 , [V] (variable) -7.47735 0.00003 (A) 1.774 1.776 1.752 

VO2 , press,Zero , [V] 

(constant) 

-7.48270 0.00001 (A) 1.257 1.260 1.244 

o VO2 , [V] (constant) 0.832784 0.00003 (A) 0.515 0.514 0.506 

o VO , press 2 , [V] (constant) -7.480301 0.00001 (A) 0.517 0.516 0.508 

CO2 X , o X CO2 , [-] 

CO2 V , [V] (varaible) 0.236576 0.01704 (A) 0.549 0.169 0.133 

VCO2 ,Zero , [V] (constant) -0.00440 0.10400 (A) 0.000 0.000 0.000 

VCO2 ,Span , [V] (constant) 1.629300 0.00046 (A) 0.135 0.139 0.127 

X CO2 ,Zero , [-] (constant) 0.0000 0.0001 (B) 

absolute 

- - - 

X CO2 ,Span , [-] (constant) 0.10100 0.00500 (B) 0.136 0.139 0.127 

CO2 RF , [-] (variable) 1.039 0.00000 (B) 0.549 0.169 0.133 

RFCO2 ,Zero , [-] (constant) 1.039 0.00000 (B) 0.000 0.000 0.000 

RFCO2 ,Span , [-] (constant) 6.209 0.00000 (B) 0.136 0.139 0.127 

o VCO 2 , [V] (constant) 0.033502 0.01309 (A) 0.001 0.001 0.001 

o RFCO2 , [-] (constant) 1.039 0.00000 (B) 0.414 0.030 0.006 

Values expected to contribute significantly to the total uncertainty are in bold. 
58

Table 6 Continued – Corrected heat release rate relative uncertainty 

Sensitivity coefficient, |si| Parameter/measurement 

[units] (type) 

Nominal 

value 

Relative 

standard 

uncertainty, 

u(xi) / xi

0.05 

MW 

0.65 

 MW 

2.70 

MW 

XCO, [-] 

VCO, [V] (variable) -0.01014 0.04726 (A) 0.066 0.004 0.001 

VCO,Zero, [V] (constant) -0.00320 0.14629 (A) 0.000 0.000 0.000 

VCO,Span, [V] (constant) 2.20140 0.00029 (A) 0.013 0.000 0.000 

XCO,Zero, [-] (constant) 0.0000 0.0001 (B) 

absolute 

- - - 

XCO,Span, [-] (constant) 0.09080 0.00500 (B) 0.013 0.000 0.000 

RFCO, [-] (variable) 0.62 0.00000 (B) 0.066 0.004 0.001 

RFCO,Zero, [-] (constant) 0.62 0.00000 (B) 0.054 0.004 0.001 

RFCO,Span, [-] (constant) 6.069 0.00000 (B) 0.013 0.000 0.000 o X H2O , [-] 

RH, [%] (variable) 45 0.04038 (B) 0.020 0.020 0.021 

Pamb, [torr] (variable) 754.10 0.00266 (B) 0.020 0.020 0.021 

Tamb, [o

C] (variable) 30.0 0.05095 (A) 0.035 0.035 0.036 

Ubdp,avg, [m/s], Tbdp,avg, [K], me & , [kg/s] 

Cbdp, [m3/K-kg] 

(constant) 

0.0698 0.05014 (B) 1.00 1.00 1.00 

bi, [torr/V] (constant) 0.10 0.00000 (B) 0.083 0.083 0.083 

Vbdp,i, [V] (variable) 1.80721 0.06231 (A) 0.082 0.082 0.082 

Vbdp,i,Zero, [V] (constant) -0.014417 0.79906 (A) 0.000 0.000 0.000 

Tbdp,i, [K] (variable) 343.0 0.00466 (A) 0.010 0.010 0.010 

ρair,ref, [kg/m3] (constant) 1.1767 0.00000 (B) 1.000 1.000 1.000 

Tref, [K] (constant) 300 0.00000 (B) 1.000 1.000 1.000 

D, [m] (constant) 1.52 0.00661 (B) 2.007 2.007 2.007 

Constants 

HC Hc Mass _ O2 (∆ ) , [MJ/kg 

O2] (constant) 

12.55 0.00159 (B) 0.981 1.000 1.000 

CO Hc Mass _ O2 (∆ ) , [MJ/kg 

O2], (constant) 

17.60 0.00057 (B) 0.019 0.000 0.000 

α, [-] (constant) 1.10 0.04364 (A) 0.003 0.019 0.079 

MWO2 , [kg/mole] 

(constant) 

32.00 0.00000 (B) 1.000 1.000 1.000 

MWair, [kg/mole] 

(constant) 

28.967 0.00000 (B) 1.000 1.000 1.000 

Values expected to contribute significantly to the total uncertainty are in bold. 
59

Table 6 Continued – Corrected heat release rate relative uncertainty 

Sensitivity coefficient, |si| Parameter/measurement 

[units] (type) 

Nominal 

value 

Relative 

standard 

uncertainty, 

u(xi) / xi

0.05 

MW 

0.65 

 MW 

2.70 

MW 

VO ,drift 2 , [V] 

O2

mV , [V/s] (constant) -2.43x10-7 0.00344 (A) 0.269 0.002 0.002 

O2 Vj , [V] (constant) 0.832475 0.00000 (A) 1115 81.99 18.10 

VCO ,drift 2 , [V] 

CO2

mV , [V/s] (constant) 9.09x10-7 0.14630 (A) 0.000 0.000 0.000 

CO2 Vj , [V] (constant) 0.034387 0.00403 (A) 0.414 0.031 0.007 

VCO,drift , [V] 

VCO m , [V/s] (constant) -7.95x10-7 0.71062 (A) 0.002 0.000 0.000 

VCO j , [V] (constant) -0.00536 0.02411 (A) 0.051 0.004 0.001 

t, [s] (variable) 0.00000 (B) 0.217 0.002 0.003 

crctd

t

q

u & 0.075 0.053 0.053 

Values expected to contribute significantly to the total uncertainty are in bold. 

At first glance, parameters with sensitivity coefficients much greater than unity are 

expected to be the largest contributors to the combined uncertainty. However these 

parameters (the oxygen analyzer voltage, O2 V , and the intercept for oxygen analyzer 

voltage drift measurement, O2 Vj ) have very small relative uncertainty. Their contribution 

is most significant at very low heat release rate, as evidenced by the increase in the 

magnitude of the sensitivity coefficient. The parameters with sensitivity coefficients near 

unity and standard relative uncertainty greater than 0.01 are the significant contributors 

across the entire range of heat release rate. Such parameters are the bi-directional probe 

flow constant, Cbdp,i, and the exhaust duct diameter, D (and the heat release rate of the 

fuel, HC

cH Mass O2 _ (∆ ) , for fuels of unknown composition). The relative uncertainty 

contribution from these two parameters combined is a little more than ±0.050 and 

essentially sets the lower limit of uncertainty of the system. In the most practical case, 

the fuel will be of unknown composition and the uncertainty of HC

cH Mass _ O2 (∆ ) will 

increase from ±0.002 to ±0.027, therefore increasing the lower limit of uncertainty to 

±0.058. The probe constant, Cbdp,i, and the duct diameter are employed in the mass flow 

rate measurement. Improvement in the uncertainty of both of these variables or an 

independent measurement of the mass flow rate measurement with a standard relative 

uncertainty less than ±0.050 is recommended. It is important to realize that the oxygen 

depletion for large fires is actually quite small due to the need to capture the plume in the 
60

exhaust duct and dilute it so that complete combustion will occur. The 3MWHRRF is 

designed for large fires. The lowest dilution rates are not optimum for small fires and 

therefore the uncertainty will be greater for the smaller fires, as evidenced by the 

increasingly large sensitivity coefficient for the oxygen sensor voltage, O2 V . 

It is recommended to apply the most comprehensive baseline correction, Method II, to 

data used for uncertainty analysis. This method of baseline correction reduces the codependency of variables that violate the rules of the uncertainty propagation as previously 

discussed. The baseline correction and associated uncertainty is a fourth innovation of 

this uncertainty analysis. 25 

Table 7 50 kW fire uncertainty comparison for the two methods of corrected HRR 

crctd

t

q

u & Baseline HRR, 

MW 

Method I correction Method II correction 

0.010 0.103 0.070 

0.052 0.135 0.075 

For large fires, 500 kW and above, the combined relative standard uncertainty computed 

for the Method I correction agrees with that of the Method II correction to within 0.005 

(therefore these results are omitted for clarity). The fact that the two methods give 

similar results means that the simpler Method I correction can be applied to routine test 

runs and the uncertainty computed. The uncertainty results are compared in Table 7 for 

very small fires and demonstrate a lesser degree of agreement. The two examples 

considered suggest that as the magnitude of the baseline drift increases the discrepancy 

between the uncertainty results increases. The combined uncertainty is larger for the 

Method I correction due to the co-variance between the correction and the measurement. 

This co-variance has a negligible effect for larger fires. 

4.4 Repeatability, Expanded Uncertainty, and Confirmation Comparison 

Repeatability 

The last component of the uncertainty analysis is the repeatability of the measurement 

system. The Large Fire Research Facility employs a natural gas burner to produce a 

known value of heat release rate. The burner operation and uncertainty are presented in 

Sec. 3.3 and Sec. 4.3.1, respectively. One cannot have a validated calorimeter without a 

high quality burner. The two facilities go hand in hand. 

 

25 When applying either the Method I or the Method II correction, the data should be centered about the 

mean time as discussed in Sec. 4.1 and 4.3.1. 
61

Table 8 Burner and calorimetry repeatability 

Burner HRR, MW Calorimetry HRR, 

MW 

Exhaust flow rate, 

m3

/min 

0.045 0.039 622.7 

0.047 0.049 641.3 

0.048 0.040 615.3 

0.043 0.045 829.2 

0.050 0.053 844.3 

0.050 0.052 805.2 

0.047 ± 0.003 

(Avg ± σ ) 0.046 ± 0.006 

0.619 0.623 673.9 

0.620 0.603 701.5 

0.621 0.620 689.6 

0.622 0.627 880.3 

0.625 0.605 895.2 

0.618 0.617 862.0 

0.621 ± 0.003 

 (Avg ± σ ) 0.616 ± 0.010 

2.610 2.706 870.9 

2.642 2.645 894.5 

2.624 2.656 874.1 

2.620 2.673 1038.3 

2.657 2.649 1056.9 

2.641 2.675 1035.9 

2.633 ± 0.017 

(Avg ± σ ) 2.667 ± 0.023 

Tests were conducted to quantify the repeatability of the measurement facility. The 

burner was run at three set points for heat release rate over two hood exhaust flow 

conditions, resulting in six separate conditions. Three repeat measurements were 

conducted at each condition over a period of two days. The conditions were ordered by 

first randomly choosing one of the two exhaust flow conditions to begin a series of tests 

for the three randomly ordered heat release rate settings and then repeating the process 

for the second exhaust flow condition. The two-step randomization was necessary in 

order to remain at an exhaust flow setting for a series of tests due to the difficulty of 

reproducing the exhaust flow conditions. However, the acquired data did not indicate 

that the exhaust flow rate has a measurable effect on the results; therefore the two sets of 

three repeats were combined. 

The standard deviation of the repeat measurements was used to compute the repeatability 

relative standard uncertainty, ur. The results of the repeat measurements together with 

their estimated uncertainties are given in Table 8. The results in the table were arrived at 

with the following analysis: 1) each heat release rate measurement in time was corrected 

for the baseline drift by algebraic subtraction, 2) the average heat release rate was 
62

computed from 300 measurements conducted at steady conditions, 3) the mean and 

standard deviation about the mean were computed for the HRR for the 6 measurements. 

Expanded Relative Uncertainty 

The combined uncertainty, uc, is computed by adding in quadrature the total uncertainty 

from the uncertainty propagation with the repeatability uncertainty: 

2 2

c t r u = u + u ( 41 ) 

Finally, the relative expanded uncertainty, Ue, is computed as 2uc. The results for the 

calibration burner and the heat release rate calorimeter are presented in Table 9. Ideally 

the repeatability (standard deviation) is an estimate of the total uncertainty, therefore the 

uncertainty as computed in Eq. (41) is a conservative estimate due to some level of 

double counting the uncertainty. The uncertainty computed from the equation modeling 

the heat release rate fails to account for sources of uncertainty not included in the model 

such as operator error, certain environmental conditions, or undetected equipment 

failures. The combined uncertainty is computed in this manner in order to address these 

possible sources of uncertainty. 

Table 9 Expanded relative uncertainty results 

crctd

e

q

U &

Nominal HRR, 

MW 

Burner Calorimeter 

0.05 0.115 0.273 

0.65 0.025 0.110 

2.70 0.027 0.107 

The calorimeter results of Table 9 are for a natural gas fire, which is a fire from a well 

characterized fuel where the uncertainty of the heat of combustion term is small. In the 

case that the fuel is not well characterized, the uncertainty of the heat of combustion term 

will increase and the increase in the expanded relative uncertainty (calorimeter output) is 

estimated to increase by 0.01 due to the addition in quadrature. 

Comparison of Calorimeter and Burner Output 

The results of the oxygen calorimetry should closely agree with the burner output. A 

comparison factor, CHRR, is determined from the ratio of the oxygen calorimetry and 

burner heat release rates. It is used as a validation and quality control factor for the 

calorimetry. When CHRR is observed to fall outside predetermined bounds, it is time to 

troubleshoot the systems, calorimetry and calibration burner. 

burner

crctd

HRR

q

q C &

& = ( 42 ) 
63

The results for CHRR versus burner output are plotted in Figure 20 and were arrived at 

with the following analysis: 1) each heat release rate measurement in time was corrected 

for the baseline drift by algebraic subtraction, 2) the mean heat release rate was computed 

from 300 measurements conducted at steady conditions, 3) the comparison factor was 

computed from the mean values. 

0.6

0.7

0.8

0.9

1.0

1.1

1.2

1.3

0.0 0.5 1.0 1.5 2.0 2.5

NG Burner

Calorimeter

C

HRR

Heat Input, MW 

Figure 20 Burner and calorimeter comparison with respect to heat release rate. 

The relative agreement between the oxygen calorimetry and the burner output is better 

than 0.98 for the entire range of heat release rates considered (0.05 to 2.70 MW). The 

bars represent the expanded relative uncertainty from Table 9. The uncertainty of the 

calorimetry is greater than the uncertainty of the burner heat release rate measurement 

and greater than the discrepancy between the calorimetry and the burner measurements. 

5 Internal Quality Control Procedures 

5.1 Certified Operator 

The custodial operator of the system should be adequately trained on all components of 

the system through thorough review of instrument manuals and consultation with the 

manufacturer’s technical representatives. The operator should be capable of training 

additional operators. The custodial operator is responsible for all Quality Control 

Procedures and scheduled maintenance. The operator should maintain a log book of all 

scheduled and unscheduled maintenance and facility modifications and upgrades. 
64

5.2 Calibration of Instruments 

The oxygen, carbon dioxide, and carbon monoxide meters must be calibrated with a zero 

gas and span gas prior to running a fire test. Because the results of the calibration are 

inputs to the data acquisition program the calibrations are performed using the data 

acquisition program. 

Oxygen Meter: 

The oxygen meter should be zeroed with nitrogen gas of 99.998 % purity (zero gas). (If 

zero gas is unavailable, 99.95 % purity will suffice.) The meter should be spanned with a 

gas mixture similar to ambient air. The oxygen content of the mixture must be 20.95 % 

and the balance must be nitrogen. Certification tickets should accompany the gas bottles 

and each minor component concentration should be certified accurate to ±1 % (from 

expanded uncertainty) of the concentration. The meter must be zeroed and spanned prior 

to each fire test where HRR is measured. The oxygen meter must be constantly powered 

up in order to establish a thermal equilibrium in the paramagnetic cell. If constant power 

to the meter is not an option, then it must be powered up at least 24 hours prior to use. 

Carbon Dioxide Meter:

The carbon dioxide meter should be zeroed with nitrogen gas of 99.998 % purity (zero 

gas). (If zero gas is unavailable, 99.95 % purity will suffice.) The meter should be 

spanned with a gas mixture of carbon dioxide, carbon monoxide, and nitrogen. The 

concentrations should be 2 % carbon dioxide, 0.1 % carbon monoxide with nitrogen 

making up the remainder. Certification tickets should accompany the gas bottles and 

each minor component concentration should be certified accurate to ±1 % (from 

expanded uncertainty) of the concentration. The meter must be zeroed and spanned prior 

to each fire test where HRR is measured. The carbon dioxide meter must be constantly 

powered up in order to establish a thermal equilibrium in the instrument. If constant 

power to the meter is not an option, then it must be powered up at least 24 h prior to use. 

Carbon Monoxide Meter:

The meter is identical to the carbon dioxide meter and the same requirements as listed 

above apply. 

Relative Humidity Meter:

The humidity meter should be set for 1 day readings of relative humidity (see THDX 

manual) and the temperature output should be in units of Celsius. The instrument should 

be calibrated every 12 months. It is recommended to return it to the factory for 

calibration for the greatest accuracy. The instrument may be calibrated in house with salt 

capsules and a controlled environment. (see THDX manual) 

Natural Gas Calorimeter:

The natural gas calorimeter should be run continuously during facility operating hours 

(five day week). Upon startup at the beginning of the each week, the instrument should 

be calibrated. Either a single point or double point calibration may be performed (see 

COSA CW95 manual). Calibration gases should be traceable to NIST primary standards 

for either gas composition or specific calorific value. Such gases are Ultra High Purity 
65

Methane (99.99 %), Specific BTU Methane (1010 BTU), and a certified mixture of 

methane (91 %) and ethane (9 %). Certification tickets should accompany the gas bottles 

and each component concentration should be certified accurate to ±1 % (from expanded 

uncertainty) of the concentration. At least once per month, a random weekly gas 

chromatograph report should be acquired from the local natural gas supplier to compare 

the two independent measurements of calorific values. 

Natural Gas Flowmeter:

The natural gas flowmeter, Instromet IRMA 15M-125, should be calibrated at the NIST 

Fluid Flow Facility every five years or more frequently if deemed necessary. 

Bi-Directional Probes and Pressure Transducers:

The bi-directional probes should be blown out with high-pressure air at the end of each 

day of testing. The high-pressure air should be applied for a minimum of 15 seconds. 

This procedure is very important when sooty fuels are burned. The pressure transducers 

should be calibrated annually. The calibration can be conducted in-house using a Dwyer 

Model 1430 point gage manometer. This device allows accurate imposition of the very 

low pressures. 

5.3 Calibration Records 

A record of all calibration certificates/reports for each instrument and calibration 

standards associated with the heat release rate measurement should be compiled in a 

single notebook and kept in the control room. An archived copy of this notebook should 

be kept by the facility supervisor. 

The following list is an example of such records: 

Carbon Dioxide and Carbon Monoxide span gas mixture certificate of analysis 

Oxygen span gas certificate of analysis 

Nitrogen zero gas certificate of analysis 

Natural gas calorimeter calibration gases certificate of analysis 

Natural gas calorimeter calibration curves 

Natural gas flowmeter calibration report 

Heat release rate comparison curves for the natural gas burner and calorimeter 

measured outputs 

Pressure transducer calibrations 

Oxygen analyzer calibration report 

Carbon Dioxide analyzer calibration report 

Carbon Monoxide analyzer calibration report 

5.4 Traps and Filters 

A new dry ice trap should be installed prior to each fire test. The traps must be cleaned 

and allowed to dry for at least 24 hours before use. 
66

The desiccant trap must be emptied and re-filled with dry desiccant (Drierite) if one 

quarter of the column appears to have changed color due to water. A desiccant trap may 

last for several fire tests. 

The particle filters upstream of the control room sampling pumps should be checked 

daily or prior to each fire test. If they are a grayish color, not bright white, then they 

should be replaced. 

The particle filter for the roof sampling pump should be checked monthly. It is important 

to keep a clean filter at this position so as to not increase the load on the sampling pump. 

5.5 Data Acquisition 

There should be a single version of the data acquisition program with a default 

configuration for basic heat release rate calorimetry. The input signals required for 

computing heat release rate should have permanent input channels that are not modified 

in any way. A list of these channels along with the description of the input and the 

channel settings should be readily visible to the data acquisition operator. In the event 

that modifications or replacement of instruments occur, the associated modifications to 

the data acquisition program must be documented in the log book. 

5.6 Data Archiving 

The data acquisition computer hard disk should be backed-up daily by NIST Central 

Computing. Fire test data should be archived on permanent media such as CD or DVD 

and properly filed for convenient retrieval. A copy of the archived data should reside 

outside of the Large Fire Research Facility. 

5.7 Confirmation of HRR Facility 

Confirmation of the complete facility using the natural gas calibration burner should be 

performed at the start and finish of a test series. For test series lasting multiple weeks 

intermediate confirmations should be performed at the start of each week. 

An extensive confirmation check may include numerous gas flow rates but more typically 

a three point check is performed. The gas flow rates (heat release rates) chosen should 

encompass the expected peak values26 for the test series of interest as well as lower levels 

(perhaps much lower levels, if a wide variety of HRR peak values is anticipated). 

 

26 The hood is nominally capable of measuring 3 MW fires using its highest flow rates. It has been used for 

brief 5 MW fires with good results. However, the upper limit depends strongly on plume entrainment rates, 

which can vary with the geometry of the fire and the plume’s height of origin below the hood. With the 

calibration burner, there is little smoke to indicate any filling of the hood and/or skirt-enclosed volume, so 

it is very difficult to know when the plume is partially spilling out. The best clue of plume spillage comes 

when the data are analyzed, if the highest HRR point measured falls appreciably below the straight line 

formed by all of the lower HRR points. 
67

The data should be reduced by first plotting HRR vs time and noting time intervals that 

satisfy the requirement of near constancy of the HRR calculated from the measured gas 

flow of the calibration burner. Extract the data for each such interval (preferably at least 

three minutes long but two minutes may be satisfactory) and compute the average and 

standard deviation of each such data interval. For the same time intervals, compute the 

average and standard deviation for the HRR (corrected for baseline drift) calculated from 

the hood system measured parameters (oxygen level, hood flow, etc.). Plot the two 

results versus each other as “Actual HRR” for the natural gas calibration burner output on 

the x-axis and “Apparent HRR” for the oxygen consumption calorimetry calculated result 

on the y-axis. Perform a linear least squares fit of the data using a graphing package or 

spreadsheet statistics. Compare the results to previous such data for the system. The 

slopes typically agree within 1 % to 2 %. 

6 Concluding Remarks 

The 3 Megawatt Heat Release Rate Facility (3MWHRRF) developed at NIST is capable 

of accurately measuring heat release rates in the range of 0.10 to 3.0 MW including brief 

peaks as high as 5 MW. The expanded uncertainty (95 % confidence interval) is 11 % of 

the heat release rate for fire sizes larger than 400 kW. The 3MWHRRF is able to 

accurately measure peak heat release rates for peaks with full-width-half-maximum equal 

to 15 s or greater. The design features that permit this response are described in this 

report. It is expected that this response should enable the facility to satisfactorily 

measure the peak heat release rate for most fuel assemblies. 

The heat release rate can be displayed in near real-time. Innovative programming was 

required to combine at least 30 signal outputs and to include corrections for 

sampling/instrumental delay and response times so that near real-time output could be 

obtained. The general design of the software is provided. 

A natural gas calibration burner with an accurate heat output and the capability of 

generating short, repeatable pulses is essential for validating the 3MWHRRF. An 

accurate flow metering valve enabled the burner output to have an expanded uncertainty 

equal to 2.7 % of the heat release rate. By use of a fast response flow controller, it is 

possible to generate square-wave pulse widths varying from 5 s to 60 s. The repeatability 

of the corrected pulse heights for the burner was typically within 3 % for the entire range 

of the peak heat release rates and pulse-widths studied. The design of the burner 

including the special flow control and flow metering capabilities are presented. 

The previously used uncertainty analyses for heat release rate do not appropriately 

account for the correlation effects in computing the oxygen difference relative to the 

background value and the effect of the drift in the gas analyzer outputs. A more 

appropriate approach to the uncertainty analysis has been utilized to include these effects 

by basing the analysis on the voltage outputs of the instruments and the use of a 

straightforward numerical method for uncertainty propagation computed with an 

electronic spreadsheet. The results of the analysis identify the major source of 
68

uncertainty as the bi-directional probes used to compute the exhaust duct volume flow 

rate. The accuracy of the exhaust duct volume flow measurement could be improved by 

calibrating the probes in a controlled flow or by implementing a tracer gas volume flow 

measurement. 

It is possible to fabricate a facility for making accurate heat release rate measurements by 

using mainly commercially available instruments. The 3MWHRRF utilizes 

commercially available equipment for the calorimeter, data acquisition/analysis system, 

and calibration burner. A general overview of the system hardware, operations and 

procedures was presented. Detailed information on specialized hardware, model numbers 

of the commercial products, and operational procedures is presented in greater detail in 

NISTIR 7052 [25]. It is thought that this information will be useful for testing 

organizations to modify their facility or to design a new heat release rate facility for large 

fires. 
69

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70

[12] J. S. Battye and et al., Shell Flow Meter Engineering Handbook, 2nd Edition 

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Guthrie, A. Maranghides, and G. W. Mulholland, The NIST 3 MW Quantitative 
71

Heat Release Rate Facility - Description and Procedures, NISTIR 7052 NIST, 

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72

Appendix A Nomenclature 

a combustion products expansion factor 

β ratio of moles of combustion products to moles of oxygen consumed 

φ oxygen depletion factor 

ρ density, kg/m3 

b pressure transducer calibration coefficient, torr/V 

C constant or comparison factor 

D duct diameter, m 

∆Hc heat of combustion, kJ/kg or kJ/m3 

j intercept of linear least squares fit 

k uncertainty coverage factor 

m slope of linear least squares fit 

m& mass loss rate or mass flow rate, kg/s 

M species molecular weight, kg/kmole 

P absolute pressure, Pa 

∆P differential pressure, Pa 

q& heat release rate, MW 

R universal gas constant, 8.314 kg m2/s2 K mol 

RF voltage range factor 

RH relative humidity, % 

s sensitivity coefficient 

t time, s 

T temperature, K 

u uncertainty 

U expanded uncertainty or velocity, m/s 

V instrument output voltage, V 

V& volume flow rate, m3

/s 

X species volume fraction 

Subscripts 

bdp bi-directional probe 

c combined 

D duct 

e exhaust or expanded 

inc incoming 

NG natural gas 

r repeatability 

t total 

Superscripts 

HC hydrocarbon 

o ambient conditions 
73

Appendix B Uncertainty Comparison Example 

B.1 Uncertainty in the Difference of Correlated Measurement Quantities 

The uncertainty in the volume fraction measurement of a gas is assumed to have two 

components, one a result of the uncertainty in the calibration gas, ucal, and the second 

arising from the electronic noise of the instrument, un. Two measurements of volume 

fraction are conducted, X1 and X2, with the same instrument, after a single calibration and 

at separate times. The algebraic difference of the measurements, ∆X, is the quantity of 

interest. Figure 1 is a graphical representation of the measurement example. The volume 

fraction of the calibration gas is XS and its standard uncertainty, ucal, is represented by 

error bars. Likewise, the standard uncertainty due to the noise is shown with error bars at 

measurements 1 and 2. Using propagation of uncertainty (i.e. “root-sum-of-squares”), 

one obtains the following expression for the uncertainty for X1: 

( ) ( ) . 2 2 1/ 2u X1 = ucal + un ( B1 ) 

0.2088

0.2090

0.2092

0.2094

0.2096

0.2098

0.2100

0.2102

2100 2200 2300 2400 2500 2600 2700 2800

Gas Volume Fraction

Time, s

X

1

X

S

X

2

∆X

Figure 1 Example of co-dependent difference measurement. 

A similar expression is obtained for the uncertainty for u(X2). Treating the two quantities 

as independent, one obtains an uncertainty in the difference in the oxygen volume 

fractions equal to √2 times the expression in Eq. B1). For the oxygen measurement, ucal

is typically a factor of 20 greater than un. This leads u(∆X) being approximately equal to 
74

√2ucal. Using this standard approach, one obtains an overestimate of the uncertainty by 

about a factor of 20. The actual uncertainty is more nearly equal to un than ucal. The 

basic reason for the overestimate is that, in the case of oxygen the two gas measurements 

are not independent because they share the same calibration. Therefore, if the calibration 

gas is low by 1 % of the specified value, it will be low for both measurements. 

To account for this lack of independence, the volume fraction measurements are 

expressed in terms of the calibration gas volume fraction, XS, the voltage of the analyzer 

corresponding to the calibration gas, VS, and the voltage corresponding to the gas volume 

fraction at measurement 1, V1, and to the voltage corresponding to the volume fraction at 

measurement 2, V2. In this case, the difference in the volume fractions is expressed as: 

( ) 1 2 V1 V2 V

X

X X X

S

S ∆ = − = − ( B2 ) 

Carrying out the uncertainty propagation based on XS, VS, V1, and V2, which are all 

independent measurements, one obtains the following expression for the uncertainty: 

1/ 2 2

2

2

1 2

2

2

1

1

1 2

1

2 2

( ) ( ) ( ) ( ) ( )

































−

− + 















− + 















+ − 













 = ∆

∆

V

u V

V V

V

V

u V

V V

V

V

u V

X

u X

X

u X

S

S

S

S ( B3 ) 

It is seen from Eq. (B3) that the uncertainty of the calibration, the first term, is only 

counted once. The remaining three terms account for the uncertainty of the noise. It is 

important to recognize that the last two terms are more heavily weighted when the 

difference in voltage (i.e. volume fraction) is small. To further illustrate the example, it 

is assumed that XS=0.21, VS=V1=1.0 V, V2=0.995 V, u(XS)=ucal, and 

u(VS)=u(V1)=u(V2)=un, ucal=20un, from which ∆X=0.00105. Computing the uncertainty 

using Eq. (B3): 

( )

1/ 2 2 2

2

2

0.005

1

0.005

1

0.21

( )















 



 



 −  +  



  + − + 

 



 = ∆

∆ n n n

cal u u u

u

X

u X

( ) ( ) ( ) (( ) ( )) ucal un un un un X

u X 4.76 2 200 1 4.76(20 ) 2 200 298 ( ) 1/ 2 2 2 2 1/ 2 2 2 2 2 = + × + ≈ + × = ∆∆

Recall that in the first example u(∆X) was equal to √2ucal=20√2un. Compared to the 

second example where u(∆X)=∆X 298un = 0.31un, it is clear that the uncertainty was 

overestimated1

 when the co-dependency of the variables was ignored. This example 

shows the importance of properly accounting for co-dependent variables and is one of the 

new aspects of this uncertainty assessment. 

 

1

 Although in this case failing to account for the correlation in the volume fractions lead to an overestimate 

of the uncertainty in their difference, in other cases an underestimate might occur. 
75

It is apparent from the above example that a straightforward application of Eq. (22) to the 

heat release rate equation (Eq. (13)) such as illustrated in the first example will not 

account for the co-dependent effects. To account for these effects, the procedure in the 

second example is followed to express the variables in terms of the most basic 

measurement inputs of instrument voltages, thermocouple temperatures, and constant 

parameters (universal, empirical and calibration). In this process, both the effect of the 

instrument span gas and zero gas is considered whereas in the example above only the 

span gas was considered. The relations for each instrument or combination of 

instruments for converting the basic measurement inputs to scientific units are presented 

in the main text.