CICIND Model Code for Steel Chimneys

CICIND Model Code for Steel Chimneys (Revision 1 – December 1999) Amend Ame ndmen mentt A – March March 2002

Commentaries and Appendices (December 2000)

Copyright CICIND 2000, 2002 ISBN 1-902998-17-0

Office of The Secretary, 14 The Chestnuts, Beechwood Park, Hemel Hempstead, Herts., HP3 0DZ, UK Tel: Te l: +44 (0)1442 211204 Fax: +44 (0)1442 256155 e-mail: [email protected] [email protected]

CICIND Model Code for Steel Chimneys REVISION 1 – DECEMBER 1999 COMMENTARIES AND APPENDICES TABLE OF CONTENTS Foreword . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .3 Commentary 1 – Glossary of commonly used words . . . . . . . . .3

Appendix 2 – Insulation, Linings and Protective Coatings . . .30 A2.1 A2 .1.. In Insu sulat latio ion n

Commentary 2 – Safety . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .6

A2.1.1 A2. 1.1 Gene General ral . . . . . . . . . . . . . . . . . . . . . . . . . . . .30 .30

Commentary 3 – Wind Load . Load . . . . . . . . . . . . . . . . . . . . . . . . . . . .8

A2.1.2. Insula Insulation tion Design Design . . . . . . . . . . . . . . . . . . . . .30 .30

C3.1. Win Wind d Speed . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .8

A2.1.3. Alumini Aluminium um Cladding Cladding . . . . . . . . . . . . . . . . . .30

C3.1.1.. BasicW C3.1.1 BasicWind ind Speed Speed . . . . . . . . . . . . . . . . . . . . . .8

A2.1.4. Minera Minerall Wool Wool or Foam Foam Insulation Insulation . . . . . . . . .31

C3.1.2.. Win C3.1.2 Wind d Maps . . . . . . . . . . . . . . . . . . . . . . . . . . .8

A2.1.5. Lined and and Multiflue Multiflue Chimneys Chimneys . . . . . . . . . . .31 .31

C3.1.3.. The Influence C3.1.3 Influence of Height . . . . . . . . . . . . . . . . .8

A2.2. A2. 2. Lin Lining ingss . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .31 .31

C3.2 The Gust Gust Factor Factor . . . . . . . . . . . . . . . . . . . . . . . . . . . . .15 .15

A2.2.1. A2. 2.1. Gene General ral . . . . . . . . . . . . . . . . . . . . . . . . . . . .31 .31

C3.3 C3. 3

Vort ortex ex Sheddi Shedding ng . . . . . . . . . . . . . . . . . . . . . . . . . . . .15 .15

A2.2.2. Design of of Separate Separate Liners Liners . . . . . . . . . . . . . . .31

C3.4 C3. 4

Movemen Mov ements ts in the the secon second d mode mode . . . . . . . . . . . . . . . . .16 .16

A2.2.3. Design of Linings Linings Attache Attached d

C3.5 C3. 5

Ovallin Oval ling g . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .18 .18

Continuously to the Shell . . . . . . . . . . . . . . .32

C3.5.1 Static effect effectss . . . . . . . . . . . . . . . . . . . . . . . . .18

A2.3. Recomm Recommended ended Start-up Start-up Procedu Procedures res . . . . . . . . . . . . . .32 .32

C3.5.2 C3. 5.2 Dyna Dynamic mic eff effect ectss . . . . . . . . . . . . . . . . . . . . . .20

A2.4. Protec Protective tive and Decorat Decorative ive Treatme Treatments nts . . . . . . . . . . . .32

C3.6 C3. 6

Interf Int erfere erence nce eff effect ectss . . . . . . . . . . . . . . . . . . . . . . . . . .21 .21

. . . . . . . . . . . . . . . . . . . . . . . . .33 Appendix 3 – Guyed Chimneys Chimneys .

Commentary 4 – Fatigue . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .22

A3.1. Guyed Chimney Chimney expansi expansion on . . . . . . . . . . . . . . . . . . . .33 .33

Commentary 5 – Openings . . . . . . . . . . . . . . . . . . . . . . . . . . . .24

A3.2. Guyed Chimney Chimney calculat calculations ions . . . . . . . . . . . . . . . . . . .33

Commentary 6 – Chemical Effects and Internal Corrosion . .26

A3.3 Guy Ropes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .33

C6.1. Chemic Chemical al Effects Effects . . . . . . . . . . . . . . . . . . . . . . . . . . . .26

Appendix 4 – Access Ladders . . . . . . . . . . . . . . . . . . . . . . . . . .34

C6.1.1 C6. 1.1.. Atta Attack ck Due Due to Sulphu Sulphurr Oxides Oxides . . . . . . . . . . .26 .26

A4.1. General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .34 .34

C6.1.2 Effec Effects ts of Flue Flue Gas Desuphu Desuphurisat risation ion . . . . . . .26

A4.2. Definiti Definitions ons . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .34

C6.1.3.. Attack Due C6.1.3 Due to Chlorine Chlorine,, Chlorides Chlorides

A4.3. Materi Materials als . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .34

and Fluorides Fluorides . . . . . . . . . . . . . . . . . . . . . . . .26 .26

A4.4. A4. 4. Fin Finish ish . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .34 .34

C6.2. Intern Internal al Corrosion Corrosion . . . . . . . . . . . . . . . . . . . . . . . . . . .26 .26

A4.5. String Stringers ers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .34

C6.3 C6. 3

A4.6. Rungs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .34

Selecti Sel ection on of mate materia rials ls . . . . . . . . . . . . . . . . . . . . . . . .26 .26

Appendix 1 – Base Plate Design . . . . . . . . . . . . . . . . . . . . . . . .28

A4.7. Safety Hoops Hoops . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .34

A1.1 A1. 1 Sim Simple ple base base plat plates es . . . . . . . . . . . . . . . . . . . . . . . . . . .28 .28

A4.8. Rest Platfor Platforms ms and Landing Landingss . . . . . . . . . . . . . . . . . . .35

A1.2 A1. 2 Bas Basee plates plates with gus gusset setss . . . . . . . . . . . . . . . . . . . . . .28

A4.9. Attachm Attachment ent to Chimney Chimney . . . . . . . . . . . . . . . . . . . . . . .35

A1.3 A1. 3 Bas Basepla eplates tes with with guss gussets ets and and compr compress ession ion ring ringss . . . . .28

A4.10.Access Hooks Hooks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .35 .35

A1.4 A1. 4 Gro Grouti uting ng . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .29 .29

DISCLAIMER This CICIND document is presented to the best of the knowledge of its members as a guide only. CICIND is not, nor are any of its members, to be held responsible for any failure alleged or proved to be due to adherence to recommendations or acceptance of information published by the association in a Model Code or in any other way. CICIND, Talacker 50, CH-8001, Zurich, Switzerland Copyright by CICIND, Zurich

CICIND Model Code – Commentaries and Appendices

FOREWORD In December 1999 the Second Edition of the Model Code for Steel Chimneys was published. This is now expanded by the publica tion of the Commentaries and Appendixes to this Model Code. The Intention of this volume is to explain the reasons behind the principles set out in the Model Code. It is divided into two parts. The Commentaries cover the theoretical derivation of the formulae and the principles used in the Model Code. The Appendices relate to more practical considerations.

COMMENTARY No. 1

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Cravat (2.19) – An upstand fixed to the roof, roofplate or cap plate to prevent the ingress of rain water (see cope hood). Also known as counter flashing. Cross-section – The section of the load bearing steel shell including the corrosion allowance. Damping device (2.20) – A device fitted to the structural shell to increase its structural damping. Doubling plate (2.21) – A plate fixed to the shell to reinforce it where increased stresses occur. Double skin chimney (2.22) – A chimney consisting of an outer load-bearing steel shell and an inner liner which carries the flue gases. Also known as a dual wall chimney.

GLOSSARY OF COMMONLY USED TERMS

Drag coefficient – see wind force coefficient

The numbers in brackets are given in figures C.1.1 and C.1.2., showing typical chimney designs.

Drain pipe (2.23) – A pipe which connects a tundish to a point outside the structural shell and used to remove condensate.

Access door (2.01) – A door for the entry of personnel or other means of inspection.

Flue – see liner

Aerodynamic stabilizer (2.03) – A device fitted to the structural shell to reduce wind excited oscillations by modifying vortex shedding Anchor bolts – See Holding down bolts Base cone (2.04) – A truncated cone incorporated immediately above the baseplate of a chimney. Baseplate (2.05) – A horizontal plate fixed to the base of a chimney. Also called a be aring plate. Base stool (2.07) – A construction comprising two vertical plates, welded to the chimney shell and to the baseplate, supporting a compression ring (2.14) through which a holding down bolt passes. Blanking off plate (2.08) – An imperforate plate fitted immediately beneath the inlet of a chimney to prevent the waste gases reaching the lower portion of the chimney. Also known as a false bottom.

Guy (2.24) – A wire rope attached at one end to a chimney and anchored at the other so as to provide tensile resistance to the lateral displacement of the chimney Guy band (2.25) – A steel section fitted around the outside of a chimney with provision for the attachment of guys. Guyed chimney (2.26) – A chimney in which not all externally applied loads (e.g. wind) are carried exclusively by the structural shell and for which guys are provided to ensure stability. Holding down bolts (2.27) – Bolts built into a concrete foundation, brick base or supporting framework to provide anchorage at the base of the chimney. Hoops – Horizontal rings forming a cage around ladders. Inlet (2.28) – A short duct fixed to the shell or baseplate of a chimney for the entry of flue gases.

Boiler mounted chimney – A chimney supported by a boiler and its foundation.

Intermediate cone (2.29) – A truncated cone incorporated in the chimney shell at an intermediate level.

Bracket (2.10) – A construction providing resistance to lateral displacement of the chimney and/or supporting part or all of the weight of the chimney.

Jointing flange (2.30) – A steel section fitted to the end of a chimney section to enable sections to be connected together.

Bracketed chimney (2.11) – A chimney in which not all external applied loads (e.g. wind) are carried exclusively by the structural shell and for which brackets, attached to an adjacent structure, are provided to ensure stability. Also known as a braced chimney. Breeching – see inlet (2.28) Cap plate (2.12) – A sloping or convex plate fitted to the top of the structural shell, covering the area between it and the liners and incorporating cravats through which the liners protrude. Cleaning door (2.13) – A door, normally at the base of the chimney, to permit the remova! of flue dust. Compression ring (2.14) – A steel plate welded to the shell which transfers the forces acting upon the chimney to the holding down bolts. Also known as a base ring. Cope band (2.15) – A steel section attached to the top of the chimney around its perimeter to give added strength and corrosion resistance at this level. Cope hood (2.16) – A hood fitted externally to the top of a liner, covering the upstand of the cap plate, to prevent the ingress of rain water. Corrosion test piece (2.17) – A fixed or removable steel plate insert, generally of lesser thickness than the shell of the chimney, in contact with the waste gases and fitted at strategic points where maximum corrosion is expected to occur. Cowl (2.18) – A conical or dished cap fitted to the top of the chimney to reduce the ingress of rain water. Also known as a rain cap.

Ladder boss – Aboss welded to the chimney shell into which an access hook or eye can be screwed to provide fixing for temporary ladders. Lateral supports (2.31) – Supports positioned at appropriate levels within the structural shell to locate the liners, allowing independent expansion of the shell. Lightning protection system – System to provide electrical continuity between the chimney and earth. Liners (2.32) – Flue gas ducts contained within the structural shell. Liner base (2.33) – A suitable support positioned at a convenient height above the baseplate of the structural steel shell to carry the weight of the liners. Lining (2.34) (see appendix No 2) – A material applied to the internal face of the chimney to prevent the flue gases contacting the inner surface of the steel shell. Multiflue chimney (2.35) – A group of two or more chimneys within a structural framework or a chimney comprising a group of two or more liners within a structural shell. Nett section – The section of the load bearing steel shell without corrosion allowance. Reinforcement – Structural shapes or plates at or near to shell aperatures to strengthen the shell. Roofplate (2.36) – A plate which follows the contour of the roof round the chimney where it passes through the roof of a building. Also known as flashing. Rungs – Horizontal bars in ladders.

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Safety system – Proprietary fall arrest system fixed to ladder rungs or beside the ladder to give a safe fixing for attachment of operatives’ safety harnesses. Self supporting chimney (2.37) – A chimney in which externally applied loads (e g. wind) are carried exclusively by the structural shell and which, together with the foundation, will remain stable under all design conditions without additional support. Splitter plate (2.38) – A vertical plate welded to the interior of the shell between two horizontally opposed inlets to divert the flow of the flue gases into a vertical direction and to inhibit the passage of flue gases from one inlet into the other.

Stayed chimney (2.40) – A chimney in which not all externally applied loads (e.g. wind) are carried exclusively by the structural shell and for which stays, connected to another structure, are provided to ensure stability. Stiffening ring – Horizontal members to prevent ovalling and to maintain the chimney shell circular during fabrication and transportation. Strakes – see aerodynamic stabilisers Stringer – Vertical member of a ladder to which the rungs are attached.

Stay (2.39) – A rigid member providing both tensile and compressive resistance to the lateral displacement of the chimney. Also known as a lateral brace.

Typical general arrangement of three types of self supporting steel chimney. The numbers are related to the text

2.03

2.08

Figure C1.1


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Structural shell (2.41) – The main external structure of th e chimney, excluding any reinforcing or flanges. Top cone (2.42) – A truncated cone or other device fitted at the top of a chimney to increase the gas exit velocity. Tundish (2.43) – A conical or sloping blanking off plate provided with facilities for drainage. Also known as a false bottom. Tuned mass damper – A form of damping device which employs a pendulum, tuned to the chimney’s natural frequency. The moving part of the pendulum is connected to the chimney by an energy absorbing device.

Vanes – See Aerodynamic stabilizers Venturi. – See Top cone Weatherhood (2.44) – A hood designed to shed rain water clear of the cravat and prevent its entry into the building. Also known as counter flashing. Wind force coefficient – The ratio between the wind pressure on the chimney and the equivalent pressure on the same area normal to the wind direction.

Typical general arrangement of guyed, stayed and bracketed chimneys. The numbers are related to the text

2.11

Figure C1.2

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The principal load is due to wind. The moment is proportional to the wind pressure, the extreme values of which follow a Fisher-Tippett Type 1 (FT1) distribution as described in reference 3.

COMMENTARY No. 2 SAFETY The safety of a chimney is ensured by the use of partial safety factors at the ultimate limit state. These partial safety factors are listed in paragraph 5.3 of the code. A chimney is thus deemed safe if the maximum stress due to the characteristic load, increased by the appropriate partial load factor, is less than the allowable stress, divided by the partial material safety factor. The level of wind load factor chosen ensures that premature failure due to low cycle fatigue, caused by wind gusts in the wind direction, can not occur.

in which the constants are the mode u and the dispersion 1/ . In temperate climates the product u.  5; other values may obtain elsewhere (see ref.2) Now, the characteristic wind is defined as having annual probability of being exceeded  0.02



It follows that the characteristic pressure q k  q 1 

Derivation Of The Partial Load Factor In The Wind Direction (Temperate Zones)

... (C2.1)

Where

The probability distribution function (pdf)



d P (x)   uexp(  u(x1))Ps1(x) dx s1

The 50-year wind pressure is x s50  1 

In(50) u

The resistance is assumed normally distributed with mean xr and standard deviation r The characteristic value is xr5%  xr  1.645r

nd  design life of structure (assumed to be 20 years for a steel chimney)

The load factor F 

Ks  a social criterion factor, given in table C2.1 

Table C2.1 - Social Criterion Factor Ks

Places of public assembly, Dams



In(50)  1.645r u

the pdf of the resistance is p r (x) 

0.05

Bridges

   1

r

2

exp 

1 2

xxr r

 2

The CDF for the wind pressure in period T years is PsT(q)  (Ps1(q))T

0.5

Towers, Masts, Offshore Structures

The effect of altering the period of exposure from 1 to T years is to

5

In order to use equation C2.1 it is necessary to estimate the value of nr. It is suggested [1, 2] that allowance be made for the number of people likely to be close to the structure at the time that maximum loading can be expected. Since maximum loading is most likely to occur under extreme wind conditions, it can be assumed that no-one will be climbing the chimney and no-one will be nearby, except through necessity. If we assume nd  20 years and Ks as 0.05 for “normal” chimneys and 0.005 for critical chimneys, acceptable probabilities can be estimated as summarised in table C2.2: Table C2.2 Typical failure probabilities for environmental economic risk Environment

nr

Chimney industrial area (“normal” chimney)

0.1

.05

103

1

.05

104

0.1

0.005

104

Chimney in urban area or hospital (“Critical chimney”)



xr  F 1

xr  1.645 r xr5%  In(50) xs50 1 u

0.005

Domestic, O ffice or Trade and Industry



then Ps1(x)  exp(exp(u(x1)))

nr  average number of people near the structure during the period of risk

Nature of structure

In(50) u

This is converted to standard measure by substituting q  x. u

The partial load factor for wind load in the wind direction is derived as follows by considering the social and economic consequences of failure or damage requiring the chimney’s repair or replacement. This involves deriving the acceptable probability of failure (P) during the chimney’s lifetime, using the following expression given in CIRIA (U.K.) Report No. 63, entitled “Rationalisation of Safety and Serviceability Factors in Structural Codes”[1] : P  104  Ks  nd / nr

This distribution has a Cumulative Distribution Function (CDF) given by P(q)  exp(exp((q  u)))

Ks

P

shift the mode from 1 to 1 

In(T) without altering the shape of u

the distribution. Hence the CDF is PsT(x)  exp(exp(u(x1)In(T))) 

The probability of failure is given by P FT   (1PsT(x))·pr(x)dx 0

Now the factor F  w · m where w is the wind load factor and m the material factor. Assuming m  1.1, then if w  1.4 PF20  8·104 if w  1.5 PF20  3·104 When failure is ductile, additional safety against collapse is derived from the chimney’s residual strength, after mobilisation of its allowable (yield) strength at one point of its periphery (i.e., at the ultimate limit state).

The probability of failure depends upon the statistical distributions of resistance and loading.

When failure is by buckling, additional safety is implicit in the relationship used between the allowable (critical buckling) strength and the yield strength of the material. This relationship includes an additional partial safety factor to ensure that the critical buckling stress is sufficiently below the lower bound of experimental curves used as a basis for the de sign (see ref. 5 ). For normal steel chimneys, this additional partial safety factor lies between 1.2 and 1.33, depending upon the diameter/ thickness ratio.

The resistance of a steel chimney may be taken as normally disributed with a coefficient of variation (ratio of standard deviation to mean value) approximately 10%.

It is, therefore, proved that wind load factors of 1.4 and 1.5, will ensure failure (collapse) probabilities of 103 and 104, required by “Normal” and “Critical” chimneys, respectively.

Chimney serving critical plant (“Critical chimney”)

It follows that safety factors should be chosen to give probabilities of failure of 103 for a “Normal” chimney and 104 for a “Critical” chimney.


References (1)

Report 63 “Rationalisation of safety and serviceability factors in structural codes” — CIRIA (U.K.), 1977

(2)

BS 8100 Part 2, British Standards Institution, 1996

(3)

Bierrum, N.R. — Letter to the Editor, CICIND REPORT Vol. 5, No. 1, 1989

(4)

ENV 1991-2-4, CEN, 1995

(5)

‘European Recommendations for steel construction’ — European Convention for Construction Steelwork (ECCS), 1978.

page 7

page 8


COMMENTARY 3 WIND LOAD At the time of publication of the revised CICIND Model Code for Steel Chimneys (1999), the wind load model currently used in ENV 1991-2-4 (eventually intended to form the basis of Eurocode 1, Part 2–4: Actions on Structures — Wind Actions) has been shown by calibration studies by CICIND and others to be unacceptable. In view of the time expected to elapse before an acceptable model for Eurocode 1 is agreed by all parties, CICIND have decided for the time being to retain the wind load model described in the 1988 version of this Model Code. A recent paper[1] has shown that this model gives safe and reasonably accurate estimates of the wind load on chimneys.

C3.1 Wind-speed As the basis for the wind-load, the hourly mean windspeed has been retained. The wind-load is calculated after estimating a turbulence intensity, by a “gust factor” method[2].

30 t (secs)

Fig. C3.2 – Relationship between windspeed and its averaging time Table C3.1 – Relationship between commonly quoted windspeeds at 10m height above grade for “open ground” situations Hourly mean

10-minute mean

Hourly mean

1.0

1.05

1.45

1.5

10-minute mean

0.95

1.0

1.4

1.45

5-second gust

0.7

0.75

1.0

1.05

3-second gust

0.65

0.7

0.95

1.0

C3.1.1. Basic wind-speed The basic wind-speed used in deriving wind-loads is the wind-speed averaged over one hour and measured at 10m above open ground at the chimney location, which has a probability of exceedence of once in 50 years. The value of the basic wind-speed for a given location should be obtained from data collected by meteorological stations. When wind speeds have been measured over periods less than 50 years, the value of the basic windspeed must be extrapolated using the Fisher-Tippett Type 1 expression for the statistical distribution of extreme values, as follows: P(V)  exp {-exp[ (V  u)]} Where: P(V)  probability of excedence of velocity V during the relevant period 1 

 slope of curve in Fig. C3.1

u

 intercept on vertical axis of curve in Fig. C3.1

For a probability of exceedence, once in 50 years, P(V) = 0.02 In some cases lower values for u and

1 are found (see lit. [3] ). 

The relationship between the wind-speed and the return period is given in figure C3.1 If the averaging time of the measurement is shorter than one hour, the hourly mean at 10m height may be determined using figure C3.2. In this figure the ratio between the hourly mean and shorter averaging periods of the wind-speed is given for various types of terrain. Table C3.1 gives a quick reference for “Open country” terrain situations.

Fig. C3.1 – Relationship between wind-speed and its return period

1

5-second gust

3-second gust

Note:- To convert “Fastest mile” windspeed to the above timeaveraged windspeeds, use the relationship (velocity  distance / time) to determine the time taken to traverse one mile. This time should then be entered in fig. C3.2.

C3.1.2 Wind Maps When no results of wind-speed measurements are available an indication of the basic wind-speed is given in the figures C3.3, C3.4, C3.5, C3.6, C3.7 and C3.8 for Europe, USA, Asia, Australia, Africa and Brazil. Some countries have not published wind velocity maps, chosing instead to specify wind pressure maps or wind velocities at specific locations. In such cases the customer should specify the wind velocity (vb) to be used in the design. The map showing isopleths for Africa is unofficial and should be used with caution.

C3.1.3. The influence of the height The increase of the wind-speed with height is in accordance with the power law: Vz  Vb · k p,z0 ·(z/10) Vb is the basic windspeed (i.e. measured at 10m above open, level terrain, without obstructions). The scale factor “k p,z0” and exponent “” depend on the terrain roughness around the chimney. The values kp,z0  1 and   0.14 have been chosen in the Model Code. This is assumed to cover the most common case when the chimney is not in the centre of cities and not at the sea shore, but somewhere in between and clear above the surrounding buildings. When structures such as buildings are being designed, it is normal to assume different values of  and k p,zo, relevant to the terrain considered. This, for instance, would give lower wind velocities in town centres than in open country. When tall structures, such as chimneys, are concerned, however, the wind velocity gradient continues to be influenced by the terrain over which it previously travelled. In some cases, the previous terrain continues to be of influence after the wind has travelled by as much as 5km over rougher terrain. In addition, the gust factor is a function of the turbulence, so that in town centres, even though the wind velocity may be less than in open country, the gust factor could be considerably higher, partially cancelling out the reduction in dynamic pressure. As a result of these considerations, it was decided to keep the Model Code simple and use just one terrain category.


Fig. C3.3 Wind speeds in m/s for Europe (10 min. mean) (note – to convert to Vb – hourly mean, divide by 1.05)

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8 5 ( 3 8 ) 9 0 ( 4 0 )

9 0 ( 4 0 )

F i g C 3 .4 – W i n d S p e e d s i n U S A

4 3 2 1 N . . . . o V t r M s I L ( s m e e a p i l n / u g o e a l s e i u s n o n e d a ) e : n t d r a s s a a i a n t i c s n o n t 3 r h n d e 3 e a o f n u t c r t l o l s o o p o ( 1 m b t u a r l a 0 i e e s o r n i m a r f t e a a t x l o n ) l h a a a i n t m , e e r b b d e g e o s i n o c v g o a i w e s t e g n d r a s o e g e e r 3 f s t a u n o t o l s w u s r , a i d n e u o r c i n e e n d c a t d . o u e n . h c T d s a n a e o o l p g l n d u w o a r t e s i m s r n t o i t v d o c u r e w s o c n n V i n o t i s t b d n o o r p d s d e i u i s r e i t v p i , s r o i e d h m n a i e e a t d s n l . d e b s l t u d y i s s . 1 n p e e . m 5 c t i . e h l i a e s l l w a p s i e n r d t w h o i n u d r

1 1 0 1 0 0 ( 4 ( 4 5 9 ) ) 1 2 1 0 3 ( 5 0 4 ( ) 5 8 ) A V G P H m i u a r u w a e e g r a i m t n r i o i c I i a s R l n a i c n S d o a s m o a

9 0 ( 4 0 )

1 4 0 ( 6 3 )

L o c a t i o n

1 5 0 ( 6 7 )

V 1 1 1 1 1 m 2 4 7 4 0 5 5 0 5 5 p h

( 5 ( 6 ( 7 ( 6 ( 4 6 5 6 5 7 ) ) ) ) )

1 4 0 ( 6 3 )

( m / s )

1 5 0 ( 6 7 ) S p e c i a l W i n d R e g i o n

1 4 0 ( 6 3 )

1 3 0 ( 5 8 ) 1 4 0 ( 6 3 )

1 3 0 ( 5 8 ) 1 2 0 ( 5 4 )

1 1 0 ( 4 9 )

1 0 0 ( 4 0 )

9 0 ( 4 0 )


Fig C3.5 – Basic windspeeds in m/s for Asia (hourly mean)

page 11

page 12

F i g 3 . 6 – B a s i c W i n d s p e e d s f o r A u s t r a l i a


D C B A

R e g i o n s

4 3 2 2 1 4 9 5

B a s ( h i o c u w r m y n l i d / s m s e p a e n e ) d V b


page 13

Fig C3.7 – Basic wind speed Vb in m/s for Africa. Isopleths shown dotted should be used with caution. For final designs local regulations should be used in all cases.

page 14


70

°

65

°

60

°

55

°

50

°

45

°

40

°

35

°

0

°

5

°

10

°

15

°

35 45

20

°

40

35

25

°

30

°

35

°

Fig. C3.8 – Windspeeds in m/s for Brazil (3-second gusts) Note – To convert to basic windspeed (hourly mean), divide by 1.5


page 15

C3.2 The gust factor The proposed method for the calculation of the bending moments in the chimney is based on the gust factor method (see lit. [4])

– The first term {Ka ·  · d 2 / mo} introduces negative aerodynamic damping

This conventional approach is:

– The second term {1  [y / (L · d )]2} gives the positive aerodynamic damping — important for large amplitudes and ensuring that the response is self-limiting.

wg (z)  G · wm (z) where: wg (z)  the load at level z G  the gust factor — a function of wind turbulence and the chimney’s natural frequency, damping and height wm (z)  the load due to the mean wind velocity An extension of this method has been proposed by B.J. Vickery (see lit. [5]) to account for the inertial response of a chimney and give more accurate values of the bending moments at levels above the base. This method has been adopted in the CICIND Model Code for Concrete Chimneys, Part (a) for the design of concrete shells, where steel reinforcement as well as shell thickness, varies often over the chimney height. In the case of steel chimneys however, which are lighter and shorter than concrete chimneys (giving a smaller inertial response) and for which there is less scope for changes of thickness with height, it was decided to use the simpler conventional method.

C3.3 Vortex shedding Large vortex-induced vibrations perpendicular to the wind direction may occur when the vortex shedding frequency coincides with a natural frequency f of the chimney. This occurs at a mean wind velocity “V” equal to the critical wind velocity “Vcr” determined by: V  Vcr  f · d / St

response induced by the chimney’s own motion. The aerodynamic parameters Ka and L incorporate the effects of the motion-induced response by means of aerodynamic damping:

For small amplitudes of up to approx. 5% of the diameter, the aerodynamic damping is described sufficiently accurately by the first term only. It can be seen that, when the structural damping  is much greater than the negative aerodynamic damping, y is quite small. As the two values converge, however, the increase in y becomes dramatic, until the self limiting amplitude is approached and increases become smaller (see Fig. C3.9). The maximum value “y” of the top deflection amplitude is calculated by multiplying the standard deviation y with a peak factor k p, i.e. y  k p · y. For small amplitudes below approx. 1–2% of the diameter, the peak-factor is approx. 4, corresponding to a stochastic type of vibration. For large amplitudes, the peak-factor is equal to about 1.5, corresponding to sinusoidal vibrations with constant amplitude. For intermediate amplitudes, the peak-factor increases gradually with decreasing amplitude. However, for the sake of simplicity, the Model Code assumes a sudden change at a value of y  4% of diameter.

y

... (C3.3.1)

in which d is the predominant chimney diameter over the top third and St is Strouhal number. Vortex-induced vibrations depend strongly on mass and damping of the chimney. The risk of large vibrations is judged by the Scruton number Sc defined as: Sc 

4 ·  ·  · mo  · d 2

... (C3.3.2)

in which  is the structural damping ratio, mo is the effective mass per unit height of the chimney as defined in the model code and  is the density of air. The risk of large vortex-induced vibrations depends on a c ombination of Scruton number and large-scale turbulence intensity of the incoming wind field. High intensity of large-scale turbulence or high Scruton numbers reduces the risk of large vortex-induced vibrations. A structure with a given Scruton number may be stable in the kind of turbulence flow normally encountered but become unstable in rare cases with low turbulence occurring under stable atmospheric stratification.

 Fig. C3.9 – Relationship between  y and Structural Damping (  ) for given values of K a, mo and d

Solving equation (C3.3.3) for the standard deviation shows that the maximum value y of the top deflection amplitude (i.e. zero to maximum) can be expressed by (see Model Code equation 7.9): y/d  kp · {c1  (c12  c2)0.5}0.5

... (C3.3.4)

where the constants c1 and c2 are equal to: c1  0.5· L2 · {1  [ · mo / (Ka ·  · d2)]}

... (C3.3.5)

or c1  0.5· L2 · {1  [Sc/(4·  · Ka)]}

C3.3.1 Structural Amplitudes The standard deviation “y” of the top structural deflection is given by, see ref. [6]: 1 d 2 d y Ca · ·  2 · St d   {Ka ·  · d 2 / mo} · { 1  [y / (L · d)]2} mo h ... (C3.3.3) in which Ca, Ka and L are aerodynamic parameters. The aerodynamic parameter Ca is found from the generalised vortexinduced wind load on structures without any significant additional

c2 

L2 ·  · d 2 · Ca2 · d Ka · mo · St4 · h

... (C3.3.6)

In smooth flow conditions, L  approximately 0.4 (see table 1), which gives the following expressions for c1 and c2 (see Model Code, 7.2.4.2): c1  0.08· {1  [ · mo / (Ka ·  · d2)]} c2 

0.16 ·  · d 3 · Ca2 Ka · mo · St4 · h

page 16


For most non-heavily damped chimneys with Scruton numbers less than 4··Ka, the influence of the constant c2 is negligible and the amplitude of the structural deflection (0 - max.) can be found from: y / d  k p · ( 2 · c1)0.5  0.4·k p · {1  [Sc/(4·  · Ka)]}0.5

... (C3.3.7)

In the present simplified and approximate approach, the aerodynamic damping parameter Ka is estimated for smooth flow cases as a function of Reynolds number (Re) only. A function of longitudinal turbulence intensity, “ I ” gives the reduction in turbulent flow, i.e.: Ka (Re, I )  Ka,max (Re) · K( I )

... (C3.3.8)

The aerodynamic damping parameter, Ka,max for smooth flow at various values of Re is given in Table 1. The function K may approximately be determined by:-

K( I )  1-31 for 0  I  0.25 and K( I )  0.25 for I  0.25. For terrain category 1 (i.e. within 5km of open sea), the minimum turbulence intensity, I min can be assumed to be 0% for wind ve locities less than or equal 10m/s and 10% for wind velocities larger than 10m/s. For all other terrain categories the minimum turbulence intensity, Imin can be assumed to be 0% for wind velocities less than or equal to 7 m/s and 10% for wind velocities larger than 7m/s. Further studies are needed to clanfy the influence of turbulence more accurately. Table C3.2. Aerodynamic parameters in smooth flow. For Reynolds numbers between the limits given, the aerodynamic parameters are determined by linear interpolation using ln(Re) as argument Re < 105

Re = 5 · 10 5

Re > 106

Ca,max

0.02

interpolation

0.01

K a,max

1.5

1.0

1.0

L

0.4

0.4

0.4

Aerodynamic parameter

Figure C3.10 shows the vortex-induced vibrations as a function of turbulence intensity for Reynolds numbers equal to 10 5 and 106, respectively.

C3.3.2 Bending Moments

The amplitude should be limited to ensure that stresses are within permissible limits, both from the point of view of failure and fatigue life. In addition, the amplitude should not be large enough to alarm bystanders. This limit is difficult to define in general terms as bystanders’ alarm is subjective, depending upon how often the response occurs, its frequency, the visibility of the chimney and the bystanders’ perception of the risk. Definition of the limiting amplitude for this aspect is, therefore left to the owner and the designer to agree for each individual case. Some guidance for highly visible chimneys with low values of Vcr ( 10m/s within 5km of sea or lake-shore,  7m/s in inland locations) is given below: Critical Chimneys – Top double amplitude (peak to peak) should be not more than 10% top diameter

The bending moments in the chimney can be calculated from the inertial load per unit length (Fw) corresponding to the relevant mode shape (i), where: Fw  m · ( 2ni)2 · i · ymax

Figure C3.10. Vortex-induced vibrations as function of turbulence intensity and Reynolds number. It is assumed that m0 / d2  50 and h/d  30, which influence the low amplitude part of the curves shown.

... (C3.3.9)

Normal Chimneys – Top double amplitude (peak to peak) should be not more than 25% top diameter These limits may be increased for less noticeable chimneys and/or those with higher values of Vcr (i.e. those which rarely see large amplitude response).

where ni  relevant natural frequency

C3.4 Movements in the second mode ymax  maximum amplitude at the relevant natural frequency or from the bending moment due to a force at 1/6 of the chimney height from the top, causing the same deflection ymax.

Just as in the case of cross-wind response in the fundamental mode, a response to excitation in the second mode, giving a top amplitude exceeding about 4% of the top diameter, triggers an increased response, initiated by the chimney’s own movement.


In the case of fundamental mode movements, response is only important to vortices shed over a length near to the chimney top, equal to about 5 top diameters, as demonstrated by Fig. C.3.11.

page 17

top amplitude in the first mode. The stresses, however, will be about the same in each case.

2.0000

[M/S r.m.s.] 2 Hz

REAL

0.0 10.000

HZ

100.00

Vcr  5.9 m/s, Scr  4.8, f  40 hz

Fig. C3.13 – Stresses and energy levels in first and second mode

This is partly demonstrated by measured values in a full-scale chimney — see fig. C.3.14. The measured values in this trace are of stresses at the base and it can be seen that many of the stress cycles in that part of the response in the second mode are much the same as those in the first mode. The second mode amplitudes were, however, only about 15% of the first mode amplitudes.

[M/S r.m.s.] 2 Hz

REAL

0.0 10.000

HZ

100.00

Vcr  4.2 m/s, Scr  18.7, f  29.5 hz Fig. C3.11 – Auto-Spectra of the anemometer signal (velocity signal), measured at Vcr in the wake of the model, measured over top half

The maximum ampltude in the second mode will occur at the top (see fig. C.3.12). The amplitude reduces to zero over a length of H / 4. This steep reduction means that the length over which vortex shedding is important will be much smaller in the case of second mode response.

s s e r t S e s a B

Second Mode First Mode

Time (secs) ƒ1 = 0.7 Hz

ƒ2 = 2.6 Hz

Fig. C3.14

The proposal for determining the top amplitudes in the second mode is given in fig. C.3.15. The stresses in both the first and second modes should be taken into account when dealing with the effects of fatigue. Fig. C3.12 – Mode shapes, first and second mode

In the second mode, the energy due to fluctuating wind pressures will be applied at the middle part of the chimney. The top amplitude of a chimney responding in the second mode will never be as great as that reached by the same chimney responding in the primary mode. This is because much more wind-induced energy would be required in the second mode. This is illustrated in Fig. C.3.13, which shows the bending moment causing the same amplitude in the second mode as in the first mode would require about 50 times more energy. On the other hand, the energy required to cause the same base stress in the second mode is almost the same as that in the first mode, even though top deflection in 2nd mode is much smaller. The proposed calculation method is based upon the assumption that more or less the same energy is applied in bending, whether the chimney is in the first or the second mode. It therefore follows that the top amplitude in the second mode would only be about 1/6 of the

Fig. C3.15 – Relationship between Scruton Number and top amplitude

page 18


C3.5 Ovalling The static as well as the dynamically fluctuating pressure causes a varying pressure over the circumference of the chimney. The varying wind pressure around a circular cylinder causes a “static” ovalling deformation of the cicle. The dynamics in the wind, including vortex shedding can cause a vibration of the circular shape, the lowest order mode and most likely to occur being that of ovalling.

horizontal sections of an unstiffened shell due to the total wind distribution, involving mainly the cos and cos2 terms (fig. C.3.18)

A major part of the stresses on horizontal sections is due to the transition from a circular shape at the base to an oval shape.

C3.5.1 Static ovalling load The distribution of the wind pressure around the circumference of the shell can be written as: p  p0 · {0.823  0.448cos  1.115cos2  0.400cos3

 0.113cos4  0.027cos5}

... (C.3.5.1)

where: p0  the wind pressure  0.5·  · v2

  Angle between wind direction and point on circumference under consideration The first term (0.823 · p0) is an overall suction and causes a small uniform tensile force on vertical cross sections of the shell. The second term (0.448 · p0 · cos) is the pressure in the wind direction (fig. C.3.16) and provides the derivation of the force coefficient (shape factor) of 0.7, to give a total load. It causes no departure from a circular cross-section.

Fig. C.3.18 – Circumferential wind pressure and deflected shape

Derivation of the increase in tensile stress is fairly straight-forward, as the maximum tensile stresses due to both beam flexure and restraint of ovalling deformation occur at the base at 180° to the wind direction (i.e. on the up-wind side). Clause 8.2 of the Model Code gives the expression:-

{tensile shell stress  tensile beam sress  (1  {6 / [(l/r)2 · (t/r)]}.

Fig. C.3.16 – Wind pressure and deflected shape due to p0cos term

The third term (1.115 · p0 ·cos — fig. C.3.17) causes ovalling.

Therefore, for t/r  0.008 and l/r  50, the increase in tensile stress  30%. This is probably unimportant in the design of chimney shells, which are usually governed by compressive stresses, but it is important in designing the base joint and holding-down bolts. The Model Code, therefore, calls for shell theory (or the above approximation) to be used for unstiffened chimneys with aspect ratio  25.

The position regarding compressive stresses is not so simple. Ref. (8) limited itself to consideration of stresses at the base, at 0° to the wind direction. Here, the compression due to beam flexure is reduced or even reversed by the shell stresses induced locally by restraint of ovalling deformation. However, increases in compressive stress are possible elsewhere. Increases in compressive stress are due to either of two effects:

Fig. C.3.17 – Wind pressure and deflected shape due to p0cos2 term

The remaining terms have little influence.

C3.5.1.1 Unstiffened shells C3.5.1.1.1 – Effect on vertical moments (stresses on horizontal sections) An analysis of the deformation and stresses in an unstiffened shell (assuming a rigidly fixed circular base) due to the ovalling load has been given elsewhere in the literature[8]. This considered stresses on

1) At the base and between values of  about 60° and 120° to the wind, the reduced compression stress due to beam flexure (function of ) has to be added to the compressive shell stress due to restraint of ovalling (function of 2) — see fig. C3.19. Significant increases in total compressive stress only occurr at relatively small values of t/r for l/r ratios less than 30 — see table C3.5.1

2) For relatively thick shells at low l/r ratios, increases of compression stress occurr on the down-wind side at 0° to the wind direction, at heights about 6 diameters above the base — see table C3.5.2. This is due to contraflexure effects, associated with restraint of ovalling, causing compressive stresses at this height.


t/r

beam stress max. MPa



beam stress

degrees

20

2.3

30

l/r

0.004

0.005

0.006

page 19

total stress

MPa

shell stress at  MPa

ratio at 

90

0.0

7.3

7.3

3.18

6.0

70

2.0

6.0

8.0

1.35

40

11.5

70

4.0

6.0

10.0

0.87

20

1.9

90

0.0

4.8

4.8

2.63

30

4.8

70

1.7

3.7

5.4

1.13

40

9.2

70

3.2

3.7

6.9

0.75

MPa

20

1.6

90

0.0

3.3

3.3

2.13

30

4.0

70

1.4

2.5

3.9

0.98

Table C3.5.1 – Max. Compression Stresses at Base of Unstiffened Chimney

t/r

l/r

0.011

0.010

0.008

0.006

max. comp. shell stress MPa

height (z)

total stress at z MPa

ratio

(x dia.)

beam stress at z MPa

20

0.9

6.2

1.3

2.2

1.64

30

0.9

6.2

8.8

9.7

1.11

20

0.9

6.2

1.5

2.3

1.57

30

0.9

6.2

8.8

9.7

1.10

40

0.9

6.2

23.1

24.0

1.04

20

0.8

6.2

1.8

2.7

1.43

30

0.8

6.2

11.0

11.8

1.03

20

0.4

7.4

1.2

1.6

1.32

30

0.4

7.8

11.0

11.4

1.03

Table C3.5.2 – Increases in compressive stress at 0o to wind (downwind side), about 6 diameters above base of an unstiffened chimney.

Oval

3.0

l

R Wind

CT CB

2.0

1.0

l

Circle

Flexure

= 20

R

= 30

Ovalling

.004

.006

.008

.01 . 011

t / R Net Compression

Max Compression

Fig. C3.20 – Increases in compressive stress over lower 6 diameters of an unstiffened chimney, due to shell effects

Total Tension Down Wind

0

°

90

°

180

°

Upwind

C.3.5.1.1.2 – Effect on horizontal moments (stresses on vertical sections) The distribution of ovalling pressure  1.115· p0 ·cos2 Fig. C3.19 – Stresses at chimney base

Therefore, combining both tables it can be seen that consideration of shell stresses leads to significant increases in compressive stresses, either at the base or at a height about 6 diameters above the base for l/r ratios  30. Guidance regarding these increases is given by fig. C3.20

Where p0 is the wind pressure, averaged over 5 seconds. Away from the ends of a long, unstiffened shell, the consequent bending moment at position  is m0, where: m0  and

1.115 · R2 · p0 · cos 2 4

m0 (max)  0.07·p 0 · d2(Nm/m)

... (C3.5.2)

page 20


(Note: 0.07 increased to 0.08 in Model Code (equation 7.11), to allow for effect of initial curvature) The associated deflection of an unstiffened shell at point  is w0, where: w0  and

1 2 · R4 ·1.115·p0 · cos 2 1 6 · E · t3

... (C3.5.2)

w0 (max)  0.06·p0 · d4 / (E · t3)

C3.5.1.2 Stiffened shells The addition of correctly sized circumferential stiffeners at the top and at the correct spacing will reduce shell stresses due to ovalling to negligible values. In considering the effect of stiffeners the following approach is used: Based upon the theory of shells[9], the deformation (w) at a distance (height) x from the stiffener is (with a small approximation) given by the following function:

This must be much less than w 0, say 1/5. Therefore, Ir must be, say, greater than 5 times (0.06 · d1.5 · t2.5). This will ensure ovalling stresses in the shell are reduced to about 20% of those in an unstiffened shell. i.e The spacing (L) of stiffening rings should be  0.56· d · (d/t)0.5 and the moment of inertia (Ir) of the stiffening ring (including participating shell (see Model Code Fig. 7.4) should be: Ir  0.3·d1.5 · t2.5 when L  0.56· d · (d/t)0.5

... (C.3.5.7)

Ir  0.3·d1.5 · t2.5 · L / 0.56· d · (d/t)0.5 when L  0.56·d·(d/t)0.5

... (C.3.5.8)

C3.5.2 Dynamic component of ovalling C.3.5.2.1 - Unstiffened shells The resonance frequency of the fundamental (ovalling) vibrations for an unstiffened cylinder is given by:

w  w0 · {1  ex/2 · [cos(x/2)sin(x/2)]} ... (C3.5.4) where:  / 2 

2 (3)0.25 · R

f 1 

· (t/R)0.5

Substituting  /2  1.52 · (t)0.5 / (R) 1.5 , the deformation of the stiffened shell becomes close to that of an unstiffened shell at a distance 1.58 · R · (R/t)0.5, or 0.56 · d · (d/t)0.5 The deformation of the shell above and below the stiffener is shown in fig. C3.21.

1 · 2



where E  A I

7.2·E·I t  0.49 · 2 · 4 ·A·R d

   

 E 

... (C3.5.9)

Young’s modulus of the shell Density of the shell Cross-section area of shell ( t m2 /m) Moment of inertia of shell about its vertical axis t3 4 ( m /m) 12

R, d and t  Radius, diameter and thickness of shell In the case of steel: f  2 5 6 0 · t / d2 w wo

Deformation with rings at distances



x = 1.32R

R t

... (C3.5.11)

The frequency of vortex shedding relevant to ovalling  2 · S t · V / d Therefore large scale resonant movemements can occur if: 2560 · t/d2  2 · S t · V / d For St  0.2, therefore, Vcr  6500 · t / d

Ring Saffener (Deformation Zero)

1·52x



R

R t

To ensure that ovalling vibrations do not occur, it is necessary to increase the moment of inertia of the shell to give a value of Vcr sufficiently high to avoid a build up of periodic excitation. Assuming that Vcr  30 m/s is high enough to achieve this, the required value of I is then given by:

Fig. C3.21 – Ovalling deformation of a cylinder with a stiff ring at x  0

It can be seen that the ovalling deformations and, therefore stresses, remain low (about 0.03w0) if the distance between stiffeners of infinitely high stiffness is smaller than 0.56 · d · (d/t)0.5. The maximum bending moment in the stiffener at this spacing is obtained after integration of the shear forces in the shell:M  0.028·p0 · d3 · (d/t)0.5 (Nm)

... (C3.5.5)

In order to be effective, the deformation of the stiffener under this moment must be much smaller than w0 — this requirement being more important than its strength. The deformation of the ring (with spacing  L) is obtained by integration of the bending moment M. The result is:

... (C3.5.12)

f  2 · S t · V c r / d 

1 2

 7.2·E·I  · A · R4

Giving: I

42 St2 Vcr2  · A · R4 · R2 7.2·E

... (C3.5.13)

For Vcr  30m/s, St  0.2,   7850 kg/m3 and E  210 · 109 N/m2, therefore I  7.4·106 · A · R2  1.8·106 · d2 · t (m4/m height) For an unstiffened shell, this means t3 / 12  1.85· 106 · d2 · t ... (C3.5.14) i.e. t/d must be  0.004, otherwise stiffening rings will be required to avoid the risk of ovalling vibrations.

C.3.5.2.2 – Stiffened shells When L 1.58 · R · (R/t)0.5: 0.19·p0 · R5.5 w · cos 2 E · Ir ·(t)0.5

... (C3.5.6)

Assuming the top of the chimney is stiffened by a ring satisfying equation (C3.5.8), ovalling vibrations can still occur at lower levels if the t/d ratio is  0.004. These vibrations are defined by:


Et3 w 12(12)

·t·



2 T2

2 2  x2 y2





2 2  x2 y2

4

page 21

Literature

E · t 4w w  R2 x4



w0

... (C3.5.15)

The solution is approximated by: w  w0 ·cos t·cos(2y/R)·cos(  · x/ L)

Substituting in equation (C3.5.12) gives: E · t2 {( /L)2  (2/R)2}4  {4 / (R2 · L4)} · 2  12 ·  {( /L)2  (2/R)2}2

... (C3.5.17)

An approximation is:

 (E / )0.5 · { 1 / [ R  ( 4 · L2) / (2 · R)]}

... (C3.5.18)

Therefore L2  ( /2)2 ·{[(R/2  · f ) · ( E / )0.5]  R2} Assuming that Vcr  30m/s is high enough to avoid oscillations and f  0.2·Vcr /R and substituting E  210·109 N/m2 and   7850 Kg/m3: L  18 · R, or 9 · d

... (C3.5.19)

From equation (C3.5.14), we have seen that the minimum value of I per unit height to avoid oscillations is: (m4 /m height)

Assuming the stiffener to provide the equivalent I of a length of shell  9 · d, Ir of stiffener (including participating shell — see Model Code, Fig. 7.4) ) must be:

C3.6

[2]

A.G. Davenport — “Wind structure and wind climate” — Seminar on Safety of Structures, Trondheim, 1977.

[3]

P.J. Rijkoort and J. Wieringa — “Extreme wind-speeds by compound Weibull analysis of exposure-corrected data” . Journal of Wind Engineering, no. 13, 1983.

[4]

A.G. Davenport — “Gust loading factors” — Proc. ASCE Journal Struct. Div., Vol. 93, No, ST 5, June, 1967.

[5]

B.J. Vickery — “Wind-induced loads on reinforced concrete chimneys” — Nat. Seminar on Tall Reinforced Concrete Chimneys, New Delhi, 1985.

[6]

S. O. Hansen — “Vortex Induced Vibrations of Line-Like Structures” — CICIND REPORT, Vol. 15, No. 1, March 1999

[7]

Shoei-Sheng Chen — “Flow-induced vibration of circular cylindrical structures”. Hemisphere Publishing Corporation 1987.

[8]

H. van Koten — “The Stress Distribution in Chimneys due to Wind Pressure” — CICIND REPORT Vol. 11, No. 2, 1995

[9]

H.van Koten — “Structural analysis of shells” — Technical University of Delft.

... (C3.5.16)

Where L  distance between stiffening rings w0  deformation of unstiffened shell   2 ·  · f f  frequency

Ir>1.75·105 · d3 · t

B.J. Vickery — “Wind loads and Design for Chimneys” — CICIND REPORT, Vol. 14, No. 2, 1998

2

Where w  deformation x  coordinate along the shell (i.e. vertical direction) y  coordinate along the circumference T  Time

I  1.85· 106 · d2 · t

[1]

... (C3.5.20)

Interference Effects

In considering the effect of aerodynamic interference by an upstream cylindrical structure on the cross-wind response of a chimney, it is generally accepted that the value of lift coefficient increases with the localised small-scale turbulence associated with wake buffetting[1]. In Reference [1], however, Vickery acknowledges in paragraph 5.2 that this does not explain the full increase in cross-wind response. He states that: “Across-wind response of the downstream structure is enhanced but the mechanism is not completely clear”. He assumes that a second contribution comes from reinforcement of the movement by buffeting at a similar frequency to that of vortex shedding by the downwind chimney. Presumably this reinforcement can be expressed by an increase in negative aerodynamic damping. Unfortunately little research data is yet available to define the way in which the increase in negative aerodynamic damping is affected by spacing, Scruton Number, or large-scale atmospheric turbulence. Therefore, for spacings between chimney and interfering structure less than 10 diameters, the Model Code merely recommends addition of structural damping to increase the chimney’s Scruton Number to more than 25. At this point it is unlikely that excessive response will be experienced. When research data is available, more definite design guidance can be given.

page 22


COMMENTARY No. 4 – FATIGUE When we consider the long term history of movement of a chimney subject to cross-wind movement in response to vortex excitation, we must take into account the following phenomena: (1) Movement is subject to a “start-up” and a “wind-down” phase at the beginning and end of each response excursion (see Fig. C4.1) (2) The stress at a point on the chimney tends to vary, reducing as the wind direction changes and its speed departs from its critical value, all due to atmospheric turbulence. The degree of reduction depends upon the level of turbulence.

Fig. C4.1 Typical trace of cross-wind oscillations Fig. C4.2 – Histograms of long term response of four full-scale chimneys

Further, in inland locations and at relatively high critical windspeeds, atmospheric turbulence is high enough to ensure that the maximum amplitude rarely occurs. This was demonstrated by a series of long term measurements (varying between 93 days and 322 days) of the response of four steel chimneys in Germany[1] — see fig. C4.2. It can be seen from these histograms that amplitudes exceeding 90% of maximum occurred only rarely, varying from about 10 cycles during 93 days at Aachen to about 100 cycles during 264 days at Cologne. The method in the Model Code takes these facts into account and develops a spectrum of response, using the Miner Rule to determine fatigue life. The Miner sum is:

M  (max / wn)k · (loge n)k 

... (4.1)

Where max  the maximum stress, per section 7.2.4 of the Model Code

Fig. C4.3 Load/cycle collectives for various values of 

To determine the number of load cycles(n), it is first necessary to know the number of occasions the wind will blow at its critical velocity (Vcr). This is determined from considerations of the probability of their occurrence — P(Vcr):

wn  the stress causing cracks after n cycles (per Wohler curve) k

 3 in the case of fatigue in steel



 a function (dependent upon Vcr) defining the shape of the load/cycle collective curve (Fig. C4.3) as follows:-

n

  max · {1  (logn / logn1)}

... (4.2)

  (Vcr / 8)1.2

... (4.3)

 Number of load cycles due to cross-wind excitation during the lifetime T

P(Vcr)  2 ·

Vcr (Vcr/ V )2 0 ·e V02

... (4.4)

Where Vo  wind velocity averaged over one year  approx. Vb(h)/4 Vb(h)  hourly mean velocity at chimney top, with exceedance probability of once in 50 years. It is assumed that the chimney responds at wind velocities between 1.1Vcr and 0.9Vcr.


Also a reduction has to be introduced to account for changes in the wind direction, so that the point of maximum stress is moved away from the point under consideration. The stress at a given point is proportional to cos2 and the total effect is roughly:-



2

( 1 / 2 ) ·

page 23

The load/cycle collective predictions over 20 years, calculated by equations (3) & (5) are shown by the dotted lines in Fig. C4.2. Because the spectrum was derived from long term measurements on relatively few chimneys, a modelling safety factor  1.4 is introduced in the expression for the Miner Number.

cos2 d  0.5

Literature

0

As a result, A2

n  3.15· 107 · T · f · 4 · 2 · 0 . 5 · 0 . 1 · A · e n  1.26· 107 · T · f · A · e A2

... (5) (see Model Code 8.5.2)

Where A  4 · Vcr /Vb(h) f  Resonance frequency

[1] W. Langer, H. Ruscheweyh & C. Verwiebe — “Untersuchungen des Querschnittverhalten von Original Stahlschornstein” — Forschungsbericht P. 230 [2] H. van Koten — “ A Calculation Method for the Cross-Wind Vibrations of Chimneys” — CICIND REPORT Vol. 14, No. 1, June 1998

page 24


placed normal to the shell {see Figures C5.2 & C5.3) and concentrated along the edge of the opening.

COMMENTARY No. 5 – OPENINGS Openings have to be strengthened to prevent local reduction of:

However, sudden ending of of the reinforcement above and below the opening can cause stress concentrations. These can treble stresses locally and lead to fatigue dama ge such as local cracks. To avoid this, in the case of openings with width greater than 40% of the chimney diameter locally, the vertical stiffeners should connect at each end with a horizontal stiffener extending around the full circumference (see fig. C5.2).

Strength Resistance against — fatigue — instability The strength of the cross-section with openings is the same as the strength of an undisturbed section if the section modulus is the same. This equality of section moduli is sufficient to fullfill the first condition of strength.

2 =

The moment of inertia of a circle with an opening subtended by the angle 2 is:

M W2



°



°

t R



I  d3  t / 8  {    sincos  [(2sin2) / (  )]} Derivation formulae for cross section properties of chimneys (both unreinforced and reinforced) and of chimneys with more than one opening at the same elevation are given in Table C5.1 If  is small then the value of I is close to that of the complete circle (0.125   d3  t). As  increases, however, the value of I drops rapidly (see Fig. C.5.1). The same holds for section modulus. To replace the lost material, reinforcing stiffeners are welded vertically to the chimney on each side of the opening. To be effective, the crosssection area (A) of each of the reinforcing stiffeners should be at least equal to A  1.25  R  t  (sin)0.5. A cross section with an opening is sensitive to the effects of buckling. This is due to the stiffness of the weakened cross-section being reduced by the possibility of the shell bending in or out at the edges of the opening. To prevent this the reinforcement stiffeners have to be

G

1

 

1

G



2

2

2

1

W1 R2t 1

W2 R2t 1

 2

 2

 

°



°

 

°

G

0

a

1

a 1



G

M W1

When the width of opening is less than 40% of the chimney’s diameter locally, it is not necessary to provide a horizontal stiffener extending around the full circumference and a more local arrangement may be used (see fig.. C5.3). Vertical reinforcement should be continued above and below the opening to a point where the added stress is unimportant. The code deems that continuing the reinforcement beyond horizontal stiffeners above and below the opening a distance at least 0.5 times the width of the opening will suffice.

G

1

G

2



1 1 G

2

1

G

e

A  2tr (   2 )

 2



G

a

G

1 =

Fig C5.1 – Reduction of inertia at openings

G a

G

I R3t



e

a

G A  2tr (   2 )  4a

1 1 G



a G

A  2tr (    ) e  rsin  / (  )

G

0

A  2tr (    )  2a e

tr2 sin  arcos tr (  )  a

I00  tr3 (    sin cos )  2ar2cos2

IGG  2tr3 (  /2sincos ) ZGG  IGG / rcos

IGG  2tr3 (  /2sincos )  4ar2cos2 ZGG  IGG / rcos

IGG  tr3 {sincos [2sin2 /(  )]} 1 Z GG  IGG / (ercos ) Z2GG  IGG / (re)

IGG  I00  Ae2 Z1GG  IGG / (ercos ) Z2GG  IGG / (re)

IG1G1  2tr3 (  /2sincos ) ZG1G1  IG1G1 /r

IG1G1  2tr3 (  /2sincos )  4ar2sin2 ZG1G1  IG1G1 /r

IG1G1  tr3 ( sincos ) ZG1G1  IG1G 1/r

IG1G1  tr3 ( sincos ) 2ar2sin2 ZG1G1  IG1G 1/r

Fig. C5.1 – Derivation formulae for section properties of chimneys with openings (a  reinforcement area)


page 25

Fig. C5.3 – Suggested detail of reinforcement for narrow openings (< 0.4D)

Fig. C5.2 – Suggested detail of reinforcement for wide openings (> 0.4D)

If the vertical height of the opening is more than twice its horizontal width, a stability check is needed. Guidance on such checks is given in the chapter on bending of plates under lateral loads in “Plates and shells”, by Timoshenko. When the duty of the chimney requires flue gas inlets whose width exceeds two-thirds of the structural shell’s diameter, a possible solution would be to provide a large number of small circular openings, giving a total area equivalent to that required. Reinforcement could then be threaded between the small holes and around the whole group, as required.

Even though it is reinforced to ensure the section complies with strength requirements, the presence of an opening can reduce locally the stiffness of the chimney and affect its natural frequencies. This reduced stiffness should therefore be taken into account when deriving the chimney’s dynamic response. This is done by taking account of the reduced local stiffness at the opening when calculating “x” for each section in equation 7.16 of the Model Code.

page 26

COMMENTARY 6 – CHEMICAL EFFECTS AND INTERNAL CORROSION C6.1. Chemical effects C6.1.1. Attack due to sulphur oxides The most common form of internal chemical attack is due to acids formed by the condensation of sulphur oxides in the flue gas. Sulphur is found in all solid and liquid fuels to varying degrees and can also be found in gaseous fuels. During the combustion process, nearly all sulphur in the fuel is oxidised to sulphur dioxide (SO 2) which can be absorbed by condensing water vapour to form sulphurous acid. A small quantity of sulphur dioxide (SO2) is further converted to sulphur trioxide (S03). The quantity depends in a complex manner upon the sulphur content of the fuel, the amount of excess air available during combustion, temperature in the combustion chamber and the presence of catalysts such as iron oxides. This small concentration of S03 (usually measured in PPM), gives rise to most of the acid corrosion problems encountered in chimneys. This is because on condensation, the S03 ions combine with water vapour to form sulphuric acid whose concentration can be as high as 85%. Condensation of these acids takes place when the temperature of the flue gas falls below their respective acid dew point temperatures (ADP), or when the flue gas comes into contact with a surface, at or below the relevant acid dew point temperature.


C6.1.3. Attack due to chlorine, chlorides and fluorides Chlorides are found in most solid fuels, including refuse and in many liquid fuels. It is also sometimes found as a pollutant in some FGD processes. Upon combustion chlorides are transformed into free chloride ions which, on contact with water vapour are transformed into hydrochloric acid. The highest condensation temperature at which hydrochloric acid has been found is 60°C. Thus, when any flue surface falls below this acid dew point, very serious corrosion will occur. This dew point is close to the water and sulphurous acid dew point. Even very small amounts of chlorides in combination with other condensed acids can cause serious corrosion problems. Hydrogen chloride, hydrogen fluoride and free chlorine in flue gases also become corrosive in their vapour stage. Stainless steels are attacked at temperatures above 320°C. Fluoride vapours are corrosive to stainless steels at temperature above 250°C.

C6.2. Internal Corrosion The internal corrosion allowances in table 8.2 of the Model Code are based upon limited exposure to condensing sulphuric acid per Fig C6.1. They are derived from the relationship between “Peak corrosion rate” and “S03 concentration” shown in figure C6.3. This, in turn, was derived from the upper bound of a family of curves which show the same relationship observed in practical situations. See lit. [2] and [3]. A safety factor of 4 has been used in arriving at the corrosion allowances.

The acid dew point temperature of sulphuric acid depends upon the concentration of S03 in the flue gas (see Fig C6.1). Provided the temperature of the flue gas and the surfaces with which it can come into contact are maintained 10°C above the acid dew point estirnated from Fig. C6.1, there is no danger of acid corrosion due to this cause. Alternatively, suitable acid resisting coatings can be applied to protect the steel. Guidance on suitable coatings and their performance is given in “CICIND Manual for Chimney Protective Coatings”. The acid dew point of sulphurous acid is about 65°C, a little above the water dew point. If the fuel is contaminated, other acids, such as hydrochloric and nitric acid can be expected to condense in the same temperature range. Thus, even if fuel and combustion processes are chosen to minimise production of S03, or if flue gases are scrubbed to remove most of the S03 and SO2, severe corrosion can be expected if the temperatures of the flue gas or the surfaces with which it can come into contact fall below 65°C, or the acid dew point temperature relevant to the reduced S03 concentration, if this is higher. Again, a safety margin is recommended of 10°C above the acid dew point temperature estimated from figure C6.1.

C.6.1.2 Effects of Flue Gas Desulphurisation (FGD) Despite the removal of most of the sulphur oxides during FGD, a severe corrosion risk remains. This is because, downstream of a scrubber, the flue gas is usually very wet and its temperature is often very low — low enough to be below the (low) value of acid dew point temperature (ADP) associated with the reduced sulphur oxide content. Fig. C6.2 shows the relationship between temperature and acid concentration to be expected and demonstrates that flue gas condensing at temperatures as low as 80°C can end up as quite concentrated acid. Also the flue gas often contains chlorides, carried over from the scrubbing materials. All steels except the very expensive high nickel alloys and titanium would deteriorate very quickly in this environment. To minimise the expense, methods have been developed to apply very thin sheets of alloy or titanium to the inner face of carbon steel or other vulnerable liners. Some organic coating materials have also been developed for this duty.

Fig. C.6.2 – Phase diagram: sulphuric acid – water vapour

C6.3 Guideline to choice of liner metallic materials Guidelines on the suitability of various metals and alloys for the range of chemical risks to be found in chimneys will be given in CICIND’s “Metallic Materials Manual” (to be published in 2001).

Literature [1]

“Desulphurisation Systems and their Effect on Operational Conditions in Chimneys”, Henseler, F., CICIND REPORT, Vol. 3, No. 2, 1987

[2]

“Influence of fuel oil characteristics and combustion conditions on the gas properties in water tube boilers” Bunz G., Diepenberg H, and Rundle A. — Jnl of the Institute of Fuel Sept 1967

[3]

“Prevention of cold end corrosion in industrial boilers” . Lech and Landowski — “Corrision” — March 1979


page 27

Fig. C6.1 – Relationship between ADP and SO 3 concentration

peak corrosion rates (micron/ 1000 hours)

SO1 concentration (ppm by vol) *) ppm

*)

= part per million (10 –6 )

Fig. C6-3 – Relationship between peak internal corrosion rates and SO3 concentration

page 28


APPENDIX 1 – DESIGN OF CHIMNEY BASE PLATES

Amendment A – March 2002

The maximum baseplate stress ( *) is given by the following expression:

This appendix is intended to give guidance on rationalising baseplate details. In the following calculations, base plate bearing stress ( *c) and maximum bolt tension (Pb*) are calculated for factored load and overturning moment. In the case of bases with a compression ring and/or gussets the values of *c and Pb* are calculated using elastic analysis as a reinforced concrete ring assuming the modular ratio of 12 [1],[2]. The area of steel bolts is taken as the thread root cross section area of the bolts. In chimneys requiring an increase in design tensile stress at the base on account of clause 8.2 of the Model Code, the value of Pb* should be factored accordingly.

*  1 . *c . l2 / tb2  f k / 1.1

... (A1.4)

where 1 is given by:

1 3.00 2.68 2.30 1.85 1.25 0.83 0.51 0.30 0.22

l/b 0 0.2 0.3 0.4 0.6 0.8 1.0 1.25 1.5

A.1.1 Simple baseplates, with no gussets or compression rings (Fig. A.1.1) and

l  the outstand of the basplate from the chimney shell b  distance between gussets

The baseplate stresses (*) on the tension side may be calculated using the method described in lit. [1]. For the particular case of l  4 · D :

*  2 · Pb* / tb2  f k / 1.1

... (A1.5)

Where 2 is obtained as follows: l/b 0.2 0.3 0.4 0.5 0.6 0.8 1.0 1.25 1.5

Fig. A1.1 – Simple Baseplate

On the compression side, the vertical shell force is distributed over a strip of width (2.l3  ts), where l3 is chosen to limit the pressure on the grout (*c) to no greater than f kg / 1.5. The maximum baseplate stress ( *)is then given by:

*  3 . * c . (l3 / tb)2  f k / 1.1 where fk

... (A1.1)

 characteristic strength of the bottom plate steel

Both equations A1.4 and A1.5 must be satisfied. The height of the gussets (h) should be sufficient to maintain acceptable shell stresses. The stress in the shell (*s) is given by the following expression:

*c  pressure on the grout ts

 thickness of shell

f kg  characteristic compressive strength of the grout On the tension side, the values of l1 and l4 should be adjusted to give vertical and rotational equilibrium. The active circumferential length of the baseplate may be taken as 3 · l2 or the bolt spacing, whichever is the lesser. The bolt tension (Pb*) then  p * · (l1  l2) / l1

... (A1.2)

Where p* is the vertical tensile force in the shell per bolt. Assuming a distribution of baseplate stress over a length of 3 · l2:

*  2 · p * / tb2  f k / 1.1 Both equations A1.1 and A1.3 have to be satisfied.

A1.2 Baseplates with Gussets (Fig. A1.2)

2 2.38 2.28 2.07 1.87 1.65 1.33 1.06 0.81 0.62

... (A1.3)

*s  w*.[( / ts)  (3 . Rs / ts2)]  f k / 1.1

... (A1.6)

Where:  and 3 are given by: No. of gussets (equally spaced) 6 12 16 20 24 28 32 40 60 80 100 and

 1.00 1.93 2.50 3.20 3.83 4.47 5.10 6.37 9.55 12.74 15.92

3 0.53 0.26 0.20 0.16 0.13 0.11 0.098 0.079 0.052 0.039 0.031

Rs  shell radius w*  the radial force on the shell per unit height of gusset at the top of the gussets, given by the following expression: w*  3 . M * / h2 Where M* is the bending moment at the base of each gusset plate due to out of balance forces under the baseplate. M*  P* . 2D

Fig. A1.2 – Baseplate with gussets

 * c . 6 . D2 . b

per gusset on the tension side per gusset on the compression side


Amendment A – March 2002

Allowance should be made for stress concentrations that may occur at the top of the gussets.

page 29

A1.4 Grouting

A1.3 Baseplate with gussets and compression ring (See Fig. A1.3)

Fig. A1.3 – Baseplate with gussets & compression ring

The baseplate stresses are calculated in the same way as in section A1.2 above using equation A1.4. The compression ring bending stresses (*) are calculated in the same way as in section A1.2 above, using equation A1.5, substituting tc (thickness of compression ring) for tb (baseplate thickness). Added to this is a direct circumferential stress arising from the out of balance moment caused by the eccentricity of the bolts, giving a total stress:

*  2 · Pb* / tc2  Pb* · N / ( 3 0 ·  · D · tc)  f K / 1.1 where N  number of bolts

Note – If the chimney is intially levelled using a nut placed on the holding down bolt under the baseplate, this nut should be loosened after packers are introduced. Fig. A1.4 provides guidance on the grouting procedure to be used under chimney baseplates.

A gusset plate thickness of 0.25D will suffice if it is of a steel whose yield strength at least equals that of the bolts.

References:

Notes regarding the derivation of 1 and 2

[1]

Brownell & Young — “Process Equipment Design”, Chapter 10

[2]

Pinfold, G.M. — “Reinforced Concrete Chimneys and Towers”

Stress coefficients 1 and 2 were obtained as follows:

1 is the coefficient applicable to the compression side and is derived from Timoshenko’s work on a rectangular plate fixed on three sides and free on the fourth. This is a reasonable assumption because pressure under the base inside the shell will produce fixity. At the gussets there is fixity by virtue of the continuity of the basplate. 2 is the coefficient applicable to the tension side. In the literature [1] this is taken from a model comprising a rectangular plate simply supported on all sides, with a patch load at th e centre representing the bearing of the nut. This is not a true reflection of the boundary conditions which are more truly fixed on two opposite sides (at the gusstes), one side being pinned (at the shell) and the fourth side free. Neither is the effect of the holding down bolt hole considered. In this Appendix, therefore, the values of 2 have been derived from plate element FE analysis, using the more realistic above boundary conditions and allowing for the bolt hole in the plate.

page 30


APPENDIX 2 – INSULATION AND PROTECTIVE LININGS AND COATINGS A.2.1

Insulation

A2.1.1 General In order to minimise loss of heat from a chimney and to maintain the temperature of the shell or liner(s) above flue gas acid dewpoint level, insulation may be fitted. But it should be appreciated that, however effective the insulation, acid will condense if the flue gas temperature entering the chimney is at or below its acid dewpoint temperature. Even if metal in contact with flue gas is generally at temperatures above its acid dewpoint, rapid local corrosion can o ccur at cold spots. In order to eliminate cold spots careful attention should be given to the following details: –

Potential air leaks should be eliminated by properly sealing flanged joints, inspection/cleaning doors, expansion joints and instrumentation apertures. The long-term effectiveness of sealing materials at the relevant service temperatures should be demonstrated.

–

Direct metal/metal contact between steel liners and the structural shell should be avoided. Liner support should incorporate a thermal isolation device.

–

Attachments such as guy ropes, aerodynamic stabilizers, ladders, platforms and pipes can act as cooling fins. Their attachment to metal in contact with flue gas should incorporate a thermal isolation device.

A2.1.2. Insulation design Insulation should be designed to maintain the surface in contact with the flue gas above acid dew point temperature everywhere, when the flue gas is at normal operating condition and at abnorrnal conditions if they can last for more than 25 hours per year (see table 7.1 of the Model Code). For design purposes, the following parameters should be used: –

Theoretical acid dewpoint, calculated taking account of sulphur content and excess combustion air should be increased by a safety margin of 10°C. If data is not available to permit calculation of the flue gas acid dew point temperature, the following values should be used for minimum metal temperature in contact with flue gas:

overall average U values W / (m2 K)

type of insulation

thickness

aluminium

6mm air gap

4.5

aluminium

18mm air gap

4.0

mineral wool

25mm

2.3*

mineral wool

50mm

1.15*

mineral wool

75mm

0.7*

mineral wool

100mm

0.5*

expanded mineral

50mm

1.15*

expanded mineral

75mm

0.7*

expanded mineral

100mm

0.5*

expanded mineral

150mm

0.35*

* These values apply for a mean insulation temperature of 40°C. They should be increased by 5% for each 50°C increase in mean insulation temperature.

Table A2.1 Typical insulation conductivities

Mineral wool or foam insulation exposed to weather should be protected by weather proofed cladding. Design of this cladding and its fixings should ensure its integrity under the action of wind at a velocity of 1.5  basic wind-speed at the relevant height (per paragraph 7.2.2.of the Model Code). The design should take account of the variation of wind pressure around the surface of the chimney at a given elevation.

A2.1.3 Aluminium cladding Aluminium cladding enclosing a narrow airspace is an effective form of insulation, due to its high thermal reflectivity. (Note — Sheet steel or other forms of cladding may be suitable in certain cases.) The exterior of the steel shell beneath the cladding should be coated with heat resisting paint. The cladding should consist of aluminium sheet not less than 1.0mm thick with symmetrical flange covers made in halves frorn aluminium sheet which also shall not be less than 1.0mm thick. The cladding should be made in strakes, using a number of equal plates per strake. All seams should be connected by aluminium alloy rivets at not more than 100mrn centres. Vertical seams of each strake should be set at the midpoint of the strake beneath.

• When fuel contains less than 0.5% by weight of sulphur, 100°C

The cladding should be fitted with its internal face the required distance away from the external face of the chimney shell, this distance being maintained by continuous circumferential spacers of the required thickness low conductivity tape coincident with the horizontal joints of the aluminium. The tape should be cemented into position by means of sodium silicate or other suitable adhesive. The ends of the horizontal rivets in the aluminium sheets serve to retain the tape in position after erection. The circumferential spacers divide the airspace between the steel shell and the aluminium cladding into sections not more than 1.5m high, thus reducing convection heat losses.

–

Ambient air temperature should be the minimum winter air temperature at the chimney location, obtained by averaging the mean temperature each night over a period of one month.

When the length of the prefabricated sections of shell between flanges is not a whole multiple of the strake width, only one make-up strake per section of chimney should be used.

–

Wind velocity should be assumed to be 5m/s.

All projections should be clad. Cleaning doors and other points where access is required should be “boxed in” with removable aluminium panels.

• When fuel is oil and/or gas, containing more than 0.5% by weight of sulphur, 175°C • When fuel is coal containing more than 0.5% by weight of sulphur, 135°C

The temperature of the metal in contact with flue gas should be checked for the condition of highest anticipated flue gas temperature. For this check the following design parameters should be assumed: –

Ambient air temperature should be maximum anticipated air temperature at the chimney location.

–

Zero wind velocity.

The design of insulation thickness to satisfy the requirements of this clause should be based upon the conductivity value of the insulation material, provided by the insulation manufacturer. If such data is not available, typical values listed in table A3.1 may be used.

The airspace at the top of the chimney should be completely sealed to prevent ingress of moisture between the steel shell and the cladding. Each upper strake of aluminium should lap over the lower strake by a minimum of 25mm, The vertical seams similarly should have a minimum lap of 25mm. To permit examination of the steel shell of the chimney without removing the cladding, 150 mm square openings, located at carefully seiected points and covered by removable panels approximately 230mm square, may be provided. Suitable positions are:


– diametrically opposite any inlet

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–

To act as insulation to maintain the flue gas temperature above its acid dew point.

–

Reduce potential for aerodynamic instability.

– approximately 1,25m from the top of the chimney Great care should be taken to ensure that dissimilar metals do not come into contact with each other. If it is essential in the design that two dissimilar metals have to be connected, a suitable non-conductive and impervious film or agent should be placed between them.

A2.1.4. Mineral wool or foam Insulation Wrapping the steel shell with a suitable grade insulation material of sufficient thickness provides more effective insulation than aluminium cladding with the usual 6mm air gap. Thicknesses of over 50mm are applied in two separate layers, the outer layer being fitted so that the vertical and horizontal joints are staggered from the joints of the inner layer. If a stiffener or flange of the chimney section projects past the outer face of the insulation, it should be wrapped with an additional layer of the same thickness for at least 75mm on ea ch side of the flange or stiffener. Insulation has to be protected from the weather, a convenient way of doing this is to cover it with metal cladding, designed as descibed above. The insulation should be fixed to the steel shell by wrapping it around so that the ends butt. It can be secured in place by steel strapping. At least two bands of strapping should be used for each strake of insulation. Insulation tends to compact and slip down the surface of the steel during transportation and erection thus leaving bare patches of steel which are potential “cold spots”. The slipping of the insulation may be prevented by welding steel pins to the shell. On low chemical load chimneys the pins can project through the insulation and have spring retaining washers fitted. On medium chemical load chimneys it is advisable to use short pins which only project half the thickne ss of the insulation so as to prevent “cold spots” forming. Usually an interval of 600mm is used between the pins.

A2.1.5. Lined and multiflue chimneys

Chimney linings may be: a) Separate liners, with a space between the liners and the outer structural shell. More than one liner may be accomodated within the structural shell, to form a “multi-flue” chimney. b) Attached continuously to the inner face of the structural shell. Such linings may be either cast against the structural shell, or be applied by spray, trowel or brush. Such linings may be: – castable refractory – solid grade diatomaceous concrete – chemical resistant coatings – fibreglass reinforced plastic (FRP)

A2.2.2 Design of separate liners A2.2.2.1. General considerations For information on the design of separate liners see the “CICIND Model Code for Concrete Chimneys, Part C - Steel Liners”. Lateral support should be provided between the liner and the structural shell as near as possible to the top of the chimney. Additional lateral supports may be required at intermediate elecations between the top of the liner and its base, depending upon considerations of stability and dynamic response, but their number should be minimised as far as possible. The lateral restraints should be designed to permit the linings to expand freely both vertically and radially. A gap between the liner and its lateral restraint(s) of between 3mm and 6mm (the larger gap being appropriate for larger diameter liners) will ensure that impact damping enhances the structural damping sufficiently to avoid problems of cross-wind oscillation in most cases.

The sapce between the outer shell and the liner of a double skin chimney can be filled with mineral wool, expanded mineral, or other suitable insulator.

The liner should be designed to resist stresses due to loads imposed by the lateral restraints, as the structural shell moves under the effect of wind or earthquake.

When expanded mineral is used as insulation, the design and fabrication of the chimney must ensure that there are no voids or openings out of which the expanded mineral can leak. A suitable drain off position must be provided at the lowest point of the expanded mineral area to ensure that the expanded mineral can be drawn off if access to the interior of the chimney shell is required.

The presence of horizontal restraints between the liner and structural shell may prevent the liner from adopting a distorted shape in response to differential expansion. As a result, bending stresses may be introduced in both the liner and the structural shell, These stresses can be very high when a single liner carries flue gases from two or more sources with different temperatures. In addition, the resulting differential liner temperature will introduce secondary thermal stresses.

Notices should be fitted to the exterior of the chimney warning that the chimney has been filled with expanded mineral. After 6 to 12 months, expanded mineral insulation compacts by about 10% thus leaving areas of the liner exposed. It is essential that this void is “topped up” with more expanded mineral and that adequate provision is left in the cap plate for topping up to take place. Sometimes a second “topping-up” is necessary after a further 12 month period.

A2.2 Protective linings A2.2.1 General Linings may be required in steel chimneys for one or more of the following purposes: –

To maximise the strength of the structural shell by keeping it cool

–

As fire protection

–

To protect an externally insulated structural shell from excessively hot flue gases. These could be generated by an operational upset or occur when an energy conservation system is by-passed.

–

Corrosion protection

A cover should be provided at the top of the structural shell to give weather protection to the airspace between liner and shell. The design of this cover should permit free expansion of the liner. Sufficient radial clearance should be incorporated to permit any relative movement, between liner and shell, that may be allowed by the lateral restraint system. In the design of this cover, special attention should be paid to the integrity of its fastenings, bearing in mind the risk of acid corrosion, stress corrosion and fatigue cracking which may be caused by aerodynamic “flutter”.

A2.2.2.2 Steel liners Unprotected steel liners should not be used in conditions of high chemical load (see table 7.1 of Model Code). In conditions of low or medium chemical load, internal corrosion allowances listed in table 8.2 of the model code may be used. In conditions of high chemical load (such as downstream of FGD), unprotected steel can be replaced by (or protected by “Wallpapered” coatings of) high nickel alloys, titanium or other metals. Guidance on choice of these materials is contained in CICIND’s “Metallic Materials Manual”, to be published in 2001. Liner supports and lateral restraints should incorporate thermal insulation so as to avoid formation of localised cold spots on the lining surfaces due to conduction of heat to the structural shell.

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Consideration should be given to the risk of fire and/or high temperature excursions described in paragraphs 7.6.1 and 7.6.2 of the model code. If the risk is significant, consideration should be given to the provision of fire protection.

when wet and when dry. Only coatings should be used that have been proved capable of retaining their protective properties in these conditions throughout the life of the chimney. Also, the chosen coating material should have expansion characteristics compatible with those of the shell throughout the relevant temperature range.

A2.2.2.3. Plastic liners Plastic and FRP liners are suitable for conditions of “high chemical load” (see table 7.1 of the Model Code), combined with low temperatures. In order to prevent material degrading, the temperature of these linings should not be allowed to exceed 100°C. Short term excursions to 150°C can be tolerated if the right type of plastic is chosen, but the life is reduced. In order to ensure liner temperature is maintained below 100°C, an automatically controlled quenching system may be installed upstream of the chimney, which is activated when the flue gas temperature exceeds 100°C.

A2.2.3 Design of linings attached continuously to the shell

The start-up procedures should follow the refractory manufacturer’s instructions. If none are available, the following procedures may be used: –

Hold gas temperature in the range of 70°C–90°C for at least 3 hours.

–

Control subsequent increases in temperature and gas flow so that no part of the liner is exposed to a gas temperature increase exceeding 50°C/hr. All parts of the lining should be exposed to gas temperature at least 75% of design temperature for at least 6 hours.

These requirements also apply to old refractory linings which have been left exposed to weather and have become soaked with water.

A2.2.3.1 General Lining or coating selection criteria and quality standards to be used during surface preparation and lining installation are detailed in the CICIND “Chimney Protective Coatings Manual”.

A2.2.3.2 Castable refractory linings (including diatomaceous concrete linings) Castable refractory should be insulating type with a minimum bulk density, after drying, of 1000kg/m3. Diatomaceous concrete should be of the “solid” grade. They should be single layer construction, installed without vapour stops. They may be cast against the inner face of the steel shell or they may be applied by a gunning process. Mixing procedures and water quantities shall follow the manufacturers’ recommendations. The minimum thickness of lining shall be 50mm. Linings 50mm to 65mm thick shall be reinforced by electric welded wire mesh. The mesh should be 50  50mm with wire of minimum diameter 2mm, or it may be 100  100mm with minimum wire diameter 3mm. The mesh should be positioned 20mm from the surface of the steel shell and should be anchored to it by steel studs, welded at 450mm spacing. Linings thicker than 65mm shall be reinforced by arc welded “V” studs, randomly orientated and at a minimum spacing of 16 per square metre. A corrosion resistant metal cap should be provided at the top of the refractory to protect its horizontal surface from the weather. Providing its surface in contact with flue gas is above acid dew point, this type of lining provides corrosion protection to the steel chimney or liner to which it is applied. Application of such a lining would convert a steel chimney, classed as being under “High chemical load” when unprotected, to a “Low chemical load” classification.

A2.2.3.3 Fibreglass reinforced plastic (FRP) linings The use of plastic and FRP for linings applied to steel chimneys is severly restricted by their tendency to separate from the steel, due to differential expansion. To minimize this problem, lining temperatures should not exceed the following values: – epoxy resins, 80°C

A2.3 Recommended start-up procedures for new castable refractory in steel chimneys or liners.

– polyesters, 60°C

It is essential that the FRP linings adhere firmly to the inside face of the chimney shell so that the surface does not crack or spall. If the acid flue gas penetrates the FRP it will attack the steel shell.

A2.2.3.4 Chemical resistant coatings Guidance on the selecion and application of chemical reisistant coatings is given in the CICIND Chimney Protective Coatings Manual. In the selection of a coating for internal use, consideration should be given to the maximum temperature to which it will be subjected, b oth

A2.4 Protective and decorative treatments Treatment selection criteria and quality standards to be used during surface preparation and coating application are detailed in the CICIND “Chimney Protective Coatings Manual”. Stainless steel is normally supplied in its mill finish condition, which is a matt, light grey. Polishing to achieve a shiny finish involves extra cost. Weathering steel, unless grit blasted, may not oxidise evenly.


APPENDIX No. 3 – GUYED CHIMNEYS A3.1. Thermal expansion effects Steel chimneys are subject to thermal expansion when the shell is heated by the flue gases and, to a small extent by strong sunlight and by large variations in ambient temperature. The vertical expansion can be considerable on tall chimneys with reasonably high flue gas temperatures, especially if they are externally insulated. For example, the vertical expansion of a steel chimney with a guy band 80m above ground level and with a shell temperature of 250°C, would be 280mm. This vertical expension expansion can greatly affect the tension in the guy wires and the consequent compressive load on the chimney shell. The stresses in guy ropes and shell should be checked under both “hot” and “cold” conditions. For instance, if the guy wires are correctly tensioned when the chimney is “cold”, the vertical expansion when the chimney goes on load will increase the tension in the guy ropes, it will also increase the vertical component in the shell plate, when it could in extreme cases produce buckling. However, if the guy wires are tensioned when the chimney is “hot”, when it goes off load the chimney will reduce in height and the guy wires will lose part of their tension. This could cause more movement under wind load than is desirable. In order to avoid these problems, a compromise initial guy rope tension under cold conditions may be necessary i.e. a tension that allows some lateral deflection of the chimney under design wind and “cold” conditions, while increasing the vertical load in the chimney by a significant but safe margin under “hot” conditions. Alternatively, if a chimney is used on a constant load 24 hours a day for long periods and maintenance resources permit, the guys can initially be correctly tensioned when the chimney is cold. When the chimney starts up and is heated to its operating temperature, the guys can be readjusted to the correct tension after the chimney has expanded. As soon as the heat load is reduced and the chimney resumes its “cold” height, however, the guys must be retensioned.

A3.2. Calculations A3.2.1 Normal conditions The guyed chimneys shall be calculated taking into the consideration second order effects. The decisive winddirections which should be taken into account are given in figure A3.1

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Mm  overturning moment produced by dead-weight or other permanent loads which may act to increase combined moment Me  overturning moment produced by permanent loads which act at all times to reduce combined moment Ma  restoring moment produced by the foundation (including guy rope anchorages) without exceeding allowable material stresses or the foundation allowable bearing pressure. In determining the support provided by the windward guy ropes, the relative stiffnesses of the chimney (acting as a cantilever) and the guy ropes, including their non-linear behaviour, should be taken into account. Many modern structural computer programs have routines for analysing guyed structures, which do this automatically. If calculations are made by hand, however, guy rope tensions should first be calculated, assuming the chimney is pinned at its base. Horizontal deflections at the rope attachment points should then be determined. The stack shell should then be analysed as a cantilever, propped by springs at the rope attachment points. The stiffness of these springs is determined by the deflections and horizontal components of tension in the ropes, previously calculated. Second order effects should be considered.

A3.2.2 Abnormal conditions The stability of the chimney should be checked at 0.1  Design Windspeed, assuming one of the guy ropes to be broken.

A3.3 Guy ropes Guy ropes should be provided in at least 3 vertical planes. T he angle between any two planes should not exceed 130°. Guy ropes should not slope more than 60° to the horizontal. Guy ropes shall be of galvanized steel wire, with steel cores, complying with ISO/R346. The wires should have a minimum tensile strength of 1450 N/mm2, A completed rope should be evenly laid and free from loose wires, disturbed strands or other irregularities and should remain in this condition when properly unwound from the reel or coil. Fittings should be of galvanized steel. Prior to erection, completed guy ropes should be greased and subjected to a tensile force amounting to 20% of their minimum breaking load for a period of 30 minutes. Guy ropes and fittings should be designed so that their minimum breaking strength exceeds 3  maximum calcuiated load, due to the sum of pretension, design wind and chimney expansion. After erection and while the chimney is cold, the guy ropes should be pretensioned so as to minimise top deflection of the chimney. The pretension may be measured by the use of a suitable instrument and should be not less than 15% nor more than 30% of the calculated maximum tension due to design wind under the hot condition. Attachments of the guy ropes should be positioned sufficiently far below the chimney top to avoid corrosive effects of the flue gases. A minimum distance of 3m is recommended.

Fig. A3.1 – Wind directions for guyed chimneys

The stability of the structure and foundation as a whole or any part of it should be investigated. Weight of anchorage should be provided such that: M  1.4 Mw  1.35Mm  0.9 Me  0.9 Ma in which: M  combined moment Mw  overturning moment produced by the design wind and imposed loads

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APPENDIX No. 4 – ACCESS LADDERS A4.1. General This section specifies the requirements for steel ladders, permanently fixed to steel chimneys, to provide means of access. They are to be fixed to the chimney in a continuous vertical length interspersed with landings and/or rest platforms as required. There may be relevant local requirements or standards which are more stringent than those detailed below and, in these cases, they must be followed. An alternative to the caged ladder system is an open ladder with a proprietary safety system, either running beside the ladder or centrally between the stringers. Rest platforms as described in A5.8 should still be incorporated at the relevant levels.

A4.2. Definitions For the purpose of this appendix the following definitions shall apply: 1) Stringers. The side members of the ladder to which the rungs are fitted. 2) Safety hoop. A bar fixed to the stringers to enclose the path of persons climbing the ladder, to prevent them falling outwards.


Stringers should, if possible, be in a continuous length, but where they are in more than one length they shall be joined by fishplates on the insides of the stringers, either welded or bolted. If bolts are used they shall be countersunk on the stringer and not less than 12mm in diameter. There shall be not less than two bolts on each side of the joint.

A4.6. Rungs Rungs shall be of round bar not less than 20mm diameter. If the bar is reduced in diameter at the ends for welding, the reduced diameter shall be 6 mm less than the diameter of the bar and there shall be a 1.5mm radius at the root of the shoulder. The rungs in a ladder or flight of ladders, shall be uniformly spaced throughout at centres of 225mm minimum to 300mm maximum. The top rung shall be on the same level as the platform which shall be extended, if necessary, to limit, to not more than 75mm, the gap between the rung and platform. Alternatively the platform may be extended to replace the top rung. Rungs shall be fitted into holes drilled in the stringers and secured by welding. Rungs shall be welded to the stringers with or without shouldering. Holes in the stringers shall be drilled to give a 1mm clearance and where shouldered rungs are used, holes shall be countersunk 1.5mm to clear the root radius (see figure A4.1).

3) Rest platform. A platform provided to enable the person climbing the ladder to rest. 4) Landing. A platform provided to enable access to part of or the whole of the circumference of the chimney.

A4.3. Materials The materials used for the construction of ladders, hoops, platforms and rest platforms shall be of carbon steel and conform to Euronorm 28–32, except those components within 3 diameters of the chimney top which, in the case of chimneys carrying flue gas with high SO 2 /SO3 content, should be of high molybdenum stainless steel (ASTM 316L or similar) or should be protected by an acid-resistant coating.

A4.4. Finish All burrs, weld-flash, sharp edges and other imperfections likely to cause injury to the hands of a person using the ladder, shall be removed and made smooth before the finishing treatment. Depending on the situation and atmospheric conditions in which the ladders are to be used, they shall be given a suitable protective finish. Hot dip galvanizing is not recommended for ladder components or connections manufactured by a cold forming process. Galvanizing may only take place after drilling, bending, sawing, etc.

A4.5. Stringers Stringers shall be of flat bar of minimum dimensions 65  10mm. The stringers shall be parallel and straight throughout the rung portion and the distance between the stringers measured from the inside faces shall not be less than 300mm and not more than 450mm. The stringers shall extend upwards, to a height of not less than 1075mm above the upper platform and shall be securely fastened at their extremities. Such extension of the stringers shall not encroach on the clear width of the platform passageway. Where, in order to step from the ladder into a landing platform, it is necessary to pass between the extended portion of the stringers, these shall be opened out from platform level to provide a clear width of 600–675mm between them at handrail level. Where access to an upper platform is from the side or front of a ladder, the ladder itself shail be extended above the platform level for a distance of not less than 1075mm or equivalent handholds shall be provided.

Fig. A4.1 – Attachment of ladder rungs to stringers

A4.7. Safety hoops If safety hoops are fitted to the ladder, the following provisions shall apply. All ladders rising 2300mm or more from a lower platform or ground level to the top rung shall be fitted with safety hoops, the spacing of which shall be uniform and at intervals not exceeding 1000mm measured along the stringer. The lowermost hoop shall be fitted to the stringers at a height of 2300  0  75mm from a lower platform or ground in order to give sufficient overhead clearance when getting on to the ladder. The uppermost hoop shall be fixed in line with any guard rail to the upper platform but in any case shall be at a height of not less than 1075mm above the level of this platform.

A4.7.1. Size of hoops Circular pattern. The width across the hoop shall be 690 to 760mm. The distance from the centre line of stringers to the inside of the back of the hoop, measured at right angles to the stringers, shall be 760 to 850mm (see figure A4.2). Rectangular pattern. The width across the hoop shall be 690 to 760mm. The distance from the centre line of stringers to the inside of the back of the hoop, measured at right angles to the stringers, shall be 690 to 760mm. The radius of the corners shall be not less than 150mm (see figure A4.2). The minimum dimensions of the hoop and strap material shall be 50  8mm. At least three vertical straps shall be fitted internally to brace the hoops; one of these straps shall be at the centre back of the hoop, and the others spaced evenly between the centre back of the hoop and the ladder stringers.


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The hooks are to be used for the temporary attachment of ladders only except as noted below. A pulley is sometimes rigged from the top of a steeplejacks’ ladder for the purpose of lifting small loads for ma intenance of the chimney. It is important that such loads shall be kept as light as possible and in no circumstance should any single load exceed 50kg. If a hook is used directly for lifting purposes, the weights of the lifting device suspended from it and of the load to be lifted should together not exceed 50kg.

A4.10.3. Materials

Fig. A4.2 – Ladder hoops

Hoops and straps shall be fixed by bolting or welding. If bolts are used they shall be countersunk, inserted from the inside of the strap or hoop and shall he not less than 12mm diameter. The assembly of hoops and straps shall be suitably braced unless secured to the stringers by double bolting, or welding.

A4.8. Rest platforms and landings When required, rest platforms shall be provided at intervals of not greater than 20m. Landing places, other than working platforms, which are provided specifically at rest platforms shall be at least 825mm square and shall have a guardrail at a height of 1075mm above the platform level with an intermediate rail and toeboards. When required, landings shall be provided at suitable levels to provide access to sampling points etc. These landings are to be adequately supported from the chimney shell and shall have a minimum width of 825mm. They are to be fitted with a guardrail at 1075mm above the platform level, with an intermediate rail and toeboards.

Hooks shall be made from steel complying with the requirements of Euronorm 25–72. In a normalised condition the steel shall have a minimum tensile strength of 430N/mm2 and a maximum tensile strength of 500N/mm2. The sockets shall be made from round steel bar complying with the requirements of Euronorm 25–72.

A4.10.4. Design The design shall be as shown in figure A4.3 for the welded hooks. The design shall be as shown in figure A4.4 for the screwed hooks and sockets. It is recommended that the screwed type of hook be used on insulated chimneys i.e., those with mineral wool or aluminium cladding as the hook does not project through the insulation. This projection could cause “cold spots” on the chimney shell. An insulating spacer should be attached to the face of the socket to minimise heat conduction between the face of the socket and the surface of the aluminium cladding.

A4.9. Attachment to chimney The ladder shall be vertical except where it follows the slope of a cone section. Stringers shall be attached to the chimney by suitable connections which shall be firmly attached to the stringers and the chimney and be sufficiently close together to make the ladder rigid throughout its length. The connections shall be of sufficient length to give a clearance of not less than 200mm behind the rungs. Suitable provision shall be made at fixing points for any differential expansion (except at platforms and landings) .

A4.10. Access hooks A4.10.1. General This section specifies requirements for hooks which are intended to provide means of access for inspection and maintenance only by steeplejacks and members of similar trades who normally fit their own ladders.

Fig. A4.3 – Welded ladder hooks

The hooks may be of two types: a) Those welded permanently to the steel shell b) Those which are screwed into sockets welded to the. shell of the steel chimney

A4.10.2. Use of access hooks The hooks shall be in a vertical line on the exterior of the structure. The use of access hooks inside chimneys exposed to corrosive gases is not recommended. The first hook should be 1.2m  50  0mm above access level. The hooks should be spaced at m ultiples of 1.5m vertical centres with a local tolerance of 50mm which will accommodate the majority of the various lengths of ladders used by steeplejacks.

Fig. A4.4 – Screwed ladder hooks and bosses

CICIND Model Code for Steel Chimneys

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