NBS BUILDING SCIENCE SERIES 148

Investigation of Construction

Failure of Reinforced

Tj^g icrete Cooling Tower at

^^^ bw Island, WV

11$ 'i- RTMENT OF COMMERCE • NATIONAL BUREAU OF STANDARDS
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NBS BUILDING SCIENCE SERIES 148

Investigation of Construction Failure

of Reinforced Concrete Cooling Tower

at Willow Island, WV

H.S. Lew

S.G. Fattal

J.R. Shaver

T.A. Reinhold

B.J. Hunt

Center for Building Technology

National Engineering Laboratory

National Bureau of Standards

Washington, DC 20234

Sponsored by:

Occupational Safety and Health Administration

U.S. Department of Labor

Washington, DC 20001

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

Issued September 1982
Library of Congress Catalog Card Number: 82-600602

National Bureau of Standards Building Science Series 148

Nat. Bur. Stand. (U.S.), Bldg. Sci. Ser. 148, 156 pages (Sept. 1982)

CODEN: BSSNBV

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1982

For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402

Price $6.50

(Add 25 percent for other than U.S. mailing)
PREFACE

An investigation of the construction failure of the natural-draft concrete

cooling tower at Willow Island, West Virginia on April 27, 1978 was carried

out by the National Bureau of Standards (NBS) at the request of the Occupational

Safety and Health Administration (OSllA) , the Department of Labor. This investigation was conducted pursuant to an interagency agreement between OSFIA and NBS.

The NBS field investigation team consisted of E. 0. Pfrang, .1. 0. Bryson,

E. Anderson, S. G. Fattal, B. J. Hunt, and H. S. Lew. Throughout the course of

this investigation, the team received full cooperation from the OSHA regional

and area offices. Assistance provided by Mr. David H. Rhone, Regional Administrator and Mr. Stanley Elliot, Area Director, is gratefully acknowledged by

the NBS team.

iii
ABSTRACT

The collapse of the natural-draft hyperbolic concrete cooling tower unit no. 2 at

the Pleasants Power Station at Willow Island, West Virginia, was investigated by

the National Bureau of Standards. The investigation included on-site inspections,

laboratory tests of construction assembly components and concrete specimens, and

analytical studies.

Based on the results of these field, laboratory, and analytical investigations, it was concluded that the most probable cause of the collapse was due to the imposition of construction loads on the shell before the concrete of lift 28 had gained

adequate strength to support these loads. The analysis of the shell indicated that

the collapse initiated at the part of the shell in lift 28 where cathead no. 4 was

located. It further showed that calculated stress resultants at several points in

that part equaled or exceeded the strength of the shell in compression, bending,

and shear. The failure of these points in that part of the shell would have propagated to cause the collapse on the entire lift 28.

Key Words: collapse; concrete; concrete strength; construction; cooling tower;

failure; hyperbolic shell; shell.

iv
TABLE OF CONTENTS

Page

PREFACE iii

ABSTRACT iv

1. INTRODUCTION 1

1.1 Background 1 1.2 Objective and Scope of the Investigation 1 1.3 Organization of the Report 2

2. COOLING TOWER CONSTRUCTION 3

2.1 Introduction 3 2.2 Physical Description of Cooling Tower Unit No. 2 at

Willow Island 3 2.3 Construction Method 4

2.4 Hoisting System 6

3. SITE INVESTIGATION 9

3.1 Introduction 9 3.2 Investigation at the Site 9

3.3 Interview Statements of Workers 12

4. LABORATORY INVESTIGATIONS 15

4.1 Introduction 15

4.2 Concrete Tests 15

4.2.1 Examination of Constituents of the Concrete 16

4.2.2 Determination of Concrete Strengths and Stiffness 17

4.3 Anchor Bolt Tests 20

4.4 Hoist Cable Tests 20

4.5 Chain Hoist Test 21

4.6 Grip-Hoist Test 21

5. EXAMINATION OF CONDITIONS OF CONSTRUCTION ASSEMBLIES AT THE

TIME OF THE COLLAPSE 23

5.1 Introduction 23

5.2 Examination of Hoisting System <, 23

5.3 Examination of Scaffolding System 26

5.4 Summary 26

V
TABLE OF CONTENTS (Cont.)

Page

6. DEFINITION OF LOADS AND SHELL ANALYSIS 27

6.1 Introduction and Background Information 11

6.2 Geometry 29

6.3 Mechanism of Load Transfer 30

6.4 Weight of Scaffolding - Load Case 2 32

6.5 Live Loads - Load Case 3 32

6.6 Weight of Cathead Gantry Assembly - Load Case 4 32

6.7 Hoisting Loads - Load Case 5 33

6.8 Other Effects , 37

6.9 Analysis of Shell 38

7. INTERPRETATION OF SHELL AI^ALYSIS 41

7.1 Introduction 41

7.2 Strength of Shell Under Combined Axial Load and

Bending Moment 41

7.3 Strength of Shell Under Shear 44

7.4 Probable Mode of Failure 45

8. SUMMARY AND CONCLUSIONS 47

9 . ACKNOWLEDGEMENTS 49

10. REFERENCES 51

TABLES AND FIGURES 3 2 APPENDICES A-1

A. Patent of Wall Construction System A-1

B. Computer Programs for Shell Analysis B-1

C. Results of Chemical Analysis of Cement C-1

D. Results of Analysis of Concrete D-1

vi
1. INTRODUCTION

1.1 BACKGROUND

Two natural-draft hyperbolic concrete cooling towers were being constructed at

the Pleasants Power Station which is located on the Ohio River at Willow Island,

West Virginia. The shell of the first unit of the two towers was completed in

August, 1977. It has a base diameter of 358 ft (109 ra) and stands 430 ft

(131 m) above the ground level. Shortly after 10 a.m. on April 27, 1978, the

top portion of the second unit which had reached a height of 166 ft (51 m)

collapsed during construction. Figure 1.1 is a photograph taken after the

failure which shows the completed tower unit no, 1 and partially completed

tower unit no. 2 in the foreground. A four-level scaffolding system which was anchored to the collapsed portion of the shell fell with it killing all 51

workers who were on the scaffold.

The shell was constructed through the use of a patented lift form technique.

Except for the lower and upper portions of the tower, the construction procedure at Willow Island utilized a scheme to place a 5-ft (1.5 ra) lift per day.

At the time of failure, 28 lifts had been completed with the most recent one having been placed the previous day. The formwork which supported the less

than one day old concrete of lift 28 had been raised into place for lift 29.

According to eyewitness accounts by workers, lift 28 began to collapse when the

third bucket of concrete was being hoisted up to the working platform. It was estimated that about 1.0 yd^ (0.8 m^) of concrete had been placed for lift 29

at that time. According to an eyewitness, the entire section of lift 28

collapsed into the tower within a few minutes.

On the day of the collapse, an inspection team from the Occupational Safety and

Health Administration (OSHA) , Department of Labor, arrived at the site and

began an investigation into the collapse. A team from the National Bureau of

Standards (NBS) joined the OSHA team two days later.

1.2 OBJECTIVE AND SCOPE OF THE INVESTIGATION

The National Bureau of Standards was requested to assist the field investigation

conducted by the Occupational Safety and Health Administration at the site of

1
the collapse and to carry out a detailed study aiming at the determination of

the most probable cause of the collapse.

In response to this request, NBS carried out field, laboratory and analytical

studies. The NBS investigators also used data obtained from onsite inspections,

OSHA case records, the patent for the forming system and drawings showing

details of the tower construction.

1.3 ORGANIZATION OF THE REPORT

The report is organized in ten chapters:

Chapter 2 presents dimensions of the tower and the material used for

construction. It also describes the construction method employed for erection

of the tower and the hoisting system used to transport construction materials

from the ground level to the top of the tower.

Chapter 3 describes observations made and measurements taken by the NBS

investigation team at the Willow Island site. It also describes typical daily

construction activities and the chain of events which took place prior to the

collapse

.

Chapter 4 presents the results of concrete tests which were used to determine

the strength and stiffness gaining characteristics of the concrete. The test

results of anchor bolts, hoist cable, chain hoist and grip-hoist are also

presented.

Chapter 5 examines possible component failures of the hoisting and scaffolding

systems with the aid of the test results presented in chapter 4 as well as data

obtained from the onsite inspection presented in chapter 3.

Chapter 6 describes criteria and rationale for defining the loads which were

acting on the shell at the time of the collapse. It also presents analysis of

the shell by means of finite element computer programs.

Chapter 7 compares the results of the shell analysis with various strength

parameters of the shell and discusses the most probable mode of failure.

Chapter 8 summarizes the findings of the investigation and presents conclusions

drawn on the basis of field, laboratory and analytical studies.

Chapter 9 acknowledges those individuals who made contributions to various

phases of the investigation and to the preparation of this report.

Chapter 10 lists the references cited in the text.

2
2. COOLING TOWER CONSTRUCTION

2.1 INTRODUCTION

This chapter describes dimensions of the tower and the materials used for

construction. It also describes the method of construction and the hoisting

system. The forrawork and scaffolding systems used for construction of the

tower were supported only by the previously completed portion of the tower.

They moved up as construction progressed with no support other than from the

partially completed structure. Because such complex self-lifting forrawork and

scaffolding systems were used, they are described in some detail.

2.2 PHYSICAL DESCRIPTION OF COOLING TOWER UNIT NO. 2 AT WILLOW ISLAND

Figure 2.1 shows the elevation and plan view of cooling tower unit no . 2 . It

may be noted that both the diameter and the shell thickness changed along the

height of the tower. At the time of failure, the tower elevation had reached

166 ft (51 m) as indicated by the dotted line on the elevation. The shell

portion of the tower was supported by 80 diagonal concrete columns of 34 inches

(0.86 m) in diameter. The shell tapered from the lintel at the base (diameter

of 342 ft, 104.2 m) to the throat (diameter of 214 ft, 85.2 m) . It then flared

out with increasing height. The exterior of the shell was divided by 96 evenly

spaced ribs (see figures 1.1 and 3.1). The panel width between two adjacent

ribs varied along the height of the tower.

In the wall section, two layers of reinforcement were provided both vertically

and horizontally. The size and spacing of steel in both directions varied with

the height of the tower. A typical cross section of the wall at lift 28 is

shown along with reinforcing and splicing details in figure 2.2. The figure

shows that, within any one lift height, all of the vertical bars were spliced

in every third panel. The horizontal bars were spliced such that no two splices

in any one lift occurred in the same vertical plane for bars on the same face.

The spacing of splices along the horizontal bars were not specified in the

drawings. All reinforcing bars were specified to meet ASTM A 615 Grade 60

requirements

.

3
Normal weight concrete was used for the tower. The mix proportions per cubic

yard of concrete wereiji/

470 lbs (214 kg) Type II (moderate heat of hydration) portland

cement

Natural sand

Natural gravel

Fly ash

Water reducing agent

Air entraining agent

1174 lbs (532 kg)

1900 lbs (862 kg)

61 lbs (27 kg)

14.1 oz (417 mL)

5.0 oz (148 mL)

Approximately 27.6 gal (104 L) of water were specified to produce 4.5 in

(114 mm) slump.

2.3 CONSTRUCTION METHOD

As mentioned in section 2.1, both the thickness and the diameter of the shell

changed along its height. To accommodate these changes as well as to provide

working surfaces for workers, a complex combination of formwork and scaffolding

system was used. A schematic drawing of the formwork-scaf folding system is

shown in figure 2.3. It is seen in this figure that the entire system was supported on previously completed portions of the tower. As construction advanced,

the system traveled up the tower with no support other than from the partially

completed structure of partially matured concrete. An interior view of a similar tower under construction is shown in figure 2.4. The four-level scaffolding system was suspended from the upper part of the tower. A detailed

description of the system is illustrated in the U.S. Patent of the system,

which is reproduced in appendix A.

The lift form and scaffolding system consisted of four basic functional

components. These were:

1) Jumpform beams

2) Jacking frames

3) Stiff back channels and formwork

4) Scaffolding

A cross section through the formwork-scaf folding system is shown in figure 2.5

in which each of the above components is identified. In the text that follows,

the function of each of these components is described and illustrated with the

aid of photographs taken at Berwick, Pennsylvania where similar hyperbolic cooling towers were under construction using basically the same construction method

used at Willow Island.

Figure 2.6 shows the manner in which jumpform beams were attached to the exterior

of the shell by hexhead bolts. Typical extruded aluminum jumpform beams which

were attached to the interior and exterior of the shell are shown in figure 2.7.

Two 10 ft (3 ra) long beams were spliced together end to end with clip plates

1/ See table 4.5(a) and (b).

4
forming a 20 foot (6 m) long section. Figure 2.8 shows the 20 ft long jumpfora

beams attached to the inside and outside of the shell by means of a number of

anchor bolts with internally threaded cone-shaped inserts and the hexhead bolts

(also see figure 6, appendix A). The assembly of the anchor bolt is shown in

figure 3.8. It may be noted in figure 2.8 that the relative positions of these

jumpform beams change in an alternating sequence for each lift of the shell.

The exterior jumpform beam differs from the interior jumpforra beam in that the

flange of the exterior jumpform beam functions as the rib mold of the shell.

The exterior flange of both the interior and exterior jumpform beams has welded

lugs to receive the reciprocating and stationary pawls for engagement of the

jacking mechanism (see figure 2.7 and figures 5 and 8, appendix A).

The second major component of the system was the jacking frame (figure 2.8).

It consisted of two structural channels which rode on both faces of the outer

flange of the jumpform beam (see figure 5, appendix A). A photograph of the

jacking frame assembly is shown in figure 2.9. Movement of the jacking frame

was restricted to one direction, i.e., parallel to the jumpform beam, by means

of a set of 12 steel wheels housed in the jacking frame.

The jacking frame also contained a hydraulic ram with a connected reciprocating

pawl and a follower pawl (item 142 and 124 in figure 5, appendix A). Wlien

jacking commenced, the reciprocating pawl was actuated by the hydraulic pressure and engaged with the jumpform beam lugs raising the jacking frame in increments equivalent to the lug spacing. The spring-loaded follower pawl ratcheted

and held the jacking frame until another hydraulic cycle could be repeated.

The third major component in the construction apparatus consisted of the

stiff back channels and formwork (figures 2.8, 2.9 and 2.10). This component

held the fresh concrete in place between each rib location. The stiffback was

a structural channel vertically supported by the jacking frame's lower steel

wheel axle (see figures 6, 7 and 8, appendix A). Adjustable joists were bolted

to the stiffbacks at the rib locations and were capable of compensating for the

shell diameter variation. The plywood form was braced by the joist-stiff back

framework. The stiffback and formwork moved integrally with the jacking frame

during the entire jacking process.

Jacking took place at both the inside and outside jumpform beams simultaneously

at all 96 rib locations. Jacking was terminated once the stiffback formwork

had cleared the previous lift and had been positioned for placement of concrete

for the next lift.

The fourth major component of the lift form-scaffolding system was a four-level

scaffold (figure 2.5). Working platforms were suspended from both the inside

and the outside jacking frames at each of the 96 rib locations. Scaffold planking and guardrails spanned between the rib locations forming the working platform. At the top level, construction materials were received via the hoisting

system, steel reinforcing was distributed for placement, and concrete was delivered around the tower by Georgia buggies.

5
Level 2 was used less in the construction process. From this level the

stiff back-formwork was accessible for adjustment. Once the formwork was set

for a new pour the use of this level became diminished.

The lower levels (levels 3 & 4) provided access to the bottom half of the

jumpform beams. The 10 ft (3 m) section of the jumpforra beam was separated

from the upper half and unbolted from the shell and delivered to the top level

for new placement. Final shell surface preparation (i.e., patching, removal of

threaded insert cones, grouting) was performed from these levels. The entire

scaffolding system moved with the jacking frame to the new elevation.

The sequence of operations for the entire formwork-scaf folding system from lift

27 to lift 29 is illustrated in figures 2.11 through 2.15. Figure 2.11 depicts

the position of the jumpform beams and the scaffolding system prior to concrete

placement for lift 27. It may be noted that the inside jumpforra beam was higher

than the outside one. Two days after the completion of lift 27, the formwork

and scaffolding systems were raised and simultaneously the lower half of the

outside jumpform beam was unbolted and lifted to a new higher position, shown

in figure 2.12. Figure 2.13 shows the final position of the entire system

prior to concrete placement for lift 28. One day after the completion of lift

28, the formwork and scaffolding systems were again raised and the lower half

of the inside jumpforra beam was lifted to a new position, as shown in figure

2.14. Figure 2.15 shows the final position of the formwork and scaffolding

systems with the inside jumpform beam extending above the outside one.

The daily preparation for concrete placement consisted of five consecutive

procedures as follows:

1) Workers removed wedging and loosened stiffbacks and formwork from previously

placed concrete. Plywood forms were removed next and cut to new desired

size to accommodate changing diameter of the tower. Forms were then placed

again in the same place after being cleaned and oiled.

2) Once all forms were replaced, jacking of the entire formwork and scaffolding

systems commenced until the next predetermined elevation was reached.

3) The lowest trailing jumpform beam was unbolted and transferred to its new

top position on top of the upper half mate of the jumpform beam. Therefore,

each pair of jumpform beams (inside and outside) at each rib location "leapfrogged" one another throughout the construction process. Figures 2.11 and

2.15 illustrate this process.

4) The formwork was wedged into its new lift position, spacing tolerances

checked, reinforcing steel placed, and concrete was then ready for

placement

.

2.4 HOISTING SYSTEM

At the Willow Island site, up to and including lift 10, construction materials

were handled by moving cranes on the ground and concrete was placed into the

formwork by a pumping process. Above this level, the materials and concrete

6
were delievered to the working platform via six ceathead gantry cranes (hereafter

referred to as catheads) powered by twin-drum hoists.

The six catheads were spaced at equal intervals around the top perimeter. Each

cathead was supported from the recently completed shell structure by four legs

which were attached to the jacking frame at two adjacent rib locations. The

catheads moved up with the lift form-scaffolding system as construction

advanced

.

4 static line, which was attached to the slide plate at the interior end of the

cathead at one end (figure 2.16) and secured to an anchor point on the ground

level at the other, guided all materials hoisted to the top working level

(figure 2.17). The anchor point location changed from a position near the wall

of the tower to the center as construction progressed upward. During hoisting

operations, the tautness of the static line was adjusted by means of a grip

hoist attached to the ground anchor point to keep the material being hoisted

from hitting the scaffold at the top of the tower.

Because of the changing shell curvature, the cathead had to be periodically

adjusted to maintain a level configuration. This was done throughout the tower

construction by adjusting the pinned telescoping outer legs and by a chain

hoist which was an integral part of the counter-static line (figure 2.16) which

also counteracted the pull exerted by the static line.

Three twin-drummed diesel hoists were spaced at equal intervals around the

tower base. Each twin-drum hoist had a single operator who served two cathead

gantries (figure 2.18). A hoist line, wound on a drum, passed through two

ground-level sheaves and traveled up along the outside face of the tower to

the outside sheave of the cathead beam (figure 2.17). It continued to the

inside sheave of the cathead beam and down to the tower center ground location

where pickups were made.

The hoist load was guided along the static line by a mechanical pulley device.

Figure 2.16 shows a bucket of concrete approaching its maximum elevation at the

inside working level. The interlink to the static line through the pulley is

also shown.

The hoist system was primarily used for lifting concrete and steel reinforcing

bars to the upper working platform on the inside of the tower. As was the case

for the cathead, various adjustments to the hoisting system were required from

lift to lift due to the changing geometry of the tower.

7

3. SITE INVESTIGATION

3.1 INTRODUCTION

During the course of the investigation, the NBS team visited the Willow Island

site on five occasions. The first visit was made on April 29, 1978, two days

after the incident, while the last visit was on August 10 and 11, 1978. This

chapter describes observations made, measurements taken of static and hoist

cables, and materials retrieved from the Willow Island site. In addition, a summary of interviews conducted by OSHA personnel with workers is presented.

These data were used in the examination of possible component failure of the

hoisting and scaffolding systems in chapter 5 and in the establishment of the

loading conditions on the tower which existed at the time of collapse in

chapter 6.

3.2 INVESTIGATION AT THE SITE

Figure 3.1 shows the exterior view of tower unit no. 2. Collapse occurred as concrete was being placed for lift 29. The top 5 ft (1.5 m) of the shell (lift

28) is missing in this photograph since the entire lift had fallen into the

center of the tower leaving a jagged edge along the top of lift 27 (figure

3.2). Evenly spaced ribs on the exterior surface of the shell and 5 ft (1.5 m)

lift lines may also be seen in figure 3.1. The exterior scaffolding-type stairs

shown in the left part of the photograph were the only means of access to

and from the top of the tower by the workers (figure 3.1).

Figures 3.3 and 3.4 show debris piled upon the floor of the tower around the

base perimeter. While the general positions of the debris inside the tower

have not been drastically changed, it is known that the debris was lifted and

moved to some extent during the rescue operations. The debris mainly consists

of the cathead gantries, scaffolding and formwork systems, safety nets, and

equipment. The concrete of lift 28 apparently pulverized upon impacting the

tower floor since no large pieces of broken concrete were found.

As can be seen in figures 3.2 and 3.5, vertical bars at the top of lift 27 were

bent toward the inside of the tower. None of the vertical bars examined at

locations around the top perimeter of the tower showed any signs of fracture.

9
Figure 3.6 shows the exterior view of a portion of the shell between catheads

no, 1 and no. 2. Dark patch marks seen in this figure were made by the cement

grout used to plug the anchor bolt holes. These marks clearly indicate that

two anchor bolts were used in lift 26 and others below. A closeup view of a typical rib in lift 27 (figure 3.7) shows that while the top anchor bolt (bolt

C in figure 2.15) was broken away, the bottom anchor bolt (the thread inserts

and the crimped rod—bolt D in figure 2.15) was intact. Careful examination of

the jumpform beams on the ground revealed that in many cases the top anchor

bolts were still attached to the flange of jumpform beams by hexhead bolts (see

figure 3.8 for the anchor bolt assembly). Further examination also revealed

that there were no tears in the flange of jumpform beams at the bolt holes nor

any fragments of the hexhead bolt in the bottom anchor bolt in lift 27. It was concluded from these observations that at the time of collapse, while the top

anchor bolts were attached to jumpform beams, the bottom hexhead bolts had been

removed

.

Localized crushing of concrete occurred surrounding the bottom anchor bolt in

lift 27 at those ribs where the legs of catheads had been supported (figure 3.9).

This indicates substantial overworking on these bolts during the previous

construction stages as a result of large movements in the cathead legs.

It was observed that the anchor bolts located at the top of lift 27 (bolt C in

figure 2.15) failed in two different modes. In many cases, the anchor bolt

together with the concrete surrounding it broke away from the shell (see figure

3.10). In other cases, the bolt itself fractured into two pieces leaving a clean tapered hole in the shell (see figure 3.11). It would not be possible to

leave such a clean hole in the shell if the cone-shaped inserts had been pulled

through the wall. This indicates some bolts fractured.

At the time of the field investigations, it was reported by workers at the site

that the collapse initiated at the location where cathead no. 4 was positioned.

Because of this reason, cathead no. 4 was carefully examined. The overall

appearance of a typical cathead gantry is shown in figure 3.12. Although all

six cathead gantry assemblies deformed severely, the component parts for each

of the cathead gantries were still intact except for the slide plates which

were separated from catheads no. 4 and 5. Figure 3.13 shows the slide plate

for cathead no. 4 which was found some distance away from the main assembly. A

large bolt to which the slide plate was attached was bent severely. A similar

deformation of the bolt was also noted at other gantries (figure 3.14). The

diameter of the sheave which was attached to the slide plate as well as others

used for the hoist cable was 12 in (305 mm).

Figure 3.15 and 3.16 show the inside and outside legs of cathead no. 4. All

four legs remained straight and showed no sign of damage. As is seen in

figure 3.16 the cathead beam was severely bent. However, close examination

showed no sign of buckling of the beam. It was concluded from this observation

that the beam was bent after the collapse had initiated.

The hoist cable used was 9/16 in (24 mm) 19 x 7 non-rotating wire rope. Of the

total of six hoist cables, one per cathead gantry, only the cables for catheads

no. 4 and no. 5 were broken. The cable for cathead no. 4 was broken at a

10
distance 408.5 ft (124.5 ra) from the axis of the drara hoist, while the cable

for cathead no. 5 was broken at the hoist end just above the dead-weight steel

ball. The remaining four hoist lines were draped over the tower following the

collapse. At both catheads no. 4 and 5 the hoist cable cut a deep groove in

the shell. At cathead no. 4, the cable cut a 4 in (100 mm) groove (figure 3.17)

and at cathead no. 5 a 5 in (125 mm) groove. The part of the cable from the

bucket to the breakage point was still attached to the bucket. Close examination showed that the concrete bucket for cathead no. 4 fell to the ground with

concrete in it. Figure 3.18 shows part of the concrete which remained inside

the bucket. On the other hand, the concrete bucket for cathead no. 5 was found

empty.

The drum hoist which served cathead no. 3 and no. 4 is shown in figure 3.19.

Subsequent to the collapse, the hoist cable for cathead no. 4 was played out

for visual observation of the condition of the cable. It was noted that the

drum hoist was operational. As may be seen from figure 3.19, the operator of

the hoist sat inside the shed which housed the hoist. In addition to audio

communication with a worker who was on the top of the tower, the operator

normally used paired visual markers on the cable to judge the position of the

material being hoisted. When the painted markers on the cable passed over the

wooden cross beams located in front of the hoist, the operator reduced the

speed of the drum so that the final positioning of the material being hoisted

could be handled safely. On the hoist cable for cathead no. 4, no such markings

were noted.

Based on the measurements and the cable profile, it was determined that the

static line was 1/2 in (13 mm) diameter 6 x 19 classification wire rope. One

end of the line was attached to the slide plate and the other end was anchored

at the ground level. As was mentioned in section 2.4, the anchor point was moved as construction progressed. It was moved to the center of the tower on April 17, 1978, on the day lift 25 was placed.

The lengths of the static line and the hoist cable for catheads no. 4 and 5 were measured. The measurements were taken by a professional surveyor for

cathead no. 4 and by NBS personnel for cathead no. 5. They were:

Cathead gantry no . 4

:

Static line length = 219.2 ft (66.8 m)

Hoist cable length = 470.4 ft (143.4 ra)

Cathead gantry no. 5:

Static line length = 221.9 ft (66.7 m)

Hoist cable length = 408.6 ft (124.5 ra)

The above lengths of the hoist cables were taken from the axis of the drum to

the top of the dead weight steel ball (figure 2.16). For the static lines, the

lengths were taken from the clevis which was attached to a concrete hopper

located near the center of the tower to pivot point G of the clevis which was attached to the slide plate of the cathead gantry (see figure 6.7).

11
3.3 INTERVIEW STATEMENTS OF WORKERS

During the course of the investigation of the collapse, many of the workers at

the Willow Island site were interviewed by OSHA personnel. Based on the interviewed statements, the work schedule of a typical day can be described as follows

:

1) At approximately 5:30 a.m., the carpenters arrived at the construction site

Immediate removal of various wedges and loosening of formwork took place

around the entire tower perimeter. This operation was executed from the

second working platform. The plywood formwork was scraped, trimmed to the new desired dimensions and oiled, and put back into place between the stiff

back channels. Simultaneously, laborers removed bolts from the lower jumpform beams. These bolts were readily accessible from the lower working

platforms 3 and 4 prior to upward jacking of the scaffolding system.

2) At approximately 7:30 a.m., the iron workers arrived at the construction

site and prepared bundles of reinforcing steel for hoisting. Adjustments

were made in the static line and the reinforcing steel was hoisted to the

top. The iron workers also distributed and placed the reinforcing steel

around the tower perimeter from the top working platform.

3) The field engineer arrived at the site at approximately 7:30 a.m., and

aligned jumpforra beams at 16 control locations.

4) Jacking of the entire scaffolding and formwork system commenced at about

8:30 a.m. Upon completion of jacking the lower jumpform beams were removed

from the shell around the entire tower perimeter and raised manually to the

top level. Formwork was then repositioned and aligned for the new pour.

5) At approximately 10:00 a.m., the first bucket of concrete was hoisted to

the top level. During the hoisting of this first bucket of concrete, the

static line tension was readjusted. The concreting operation began at

catheads no. 4 and 5 and progressed in two directions toward catheads

no. 1 and 6 where the operation terminated. While concrete was being

placed, tying of reinforcing steel, jumpforra beam repositioning, surface

preparation at the cold joint and grouting, curing agent application, etc.,

took place and terrainated prior to completion of concrete placement. The

remainder of the day, until approximately 2:00 or 3:00 p.m., was used for

concrete placement for the entire five-foot lift. Following placement of

the concrete, all workmen retired and the process was repeated on the

following day.

Just prior to the collapse, the following chain of events were described by

those who saw the collapse from the center of the tower.

1) Shortly before the collapse, the first bucket of concrete was delivered

to cathead no. 4 and emptied into Georgia buggies.

2) Cathead no. 5 received its first bucket of concrete which was also emptied

into Georgia buggies.

12
When the workers at the center of the tower heard a loud cracking sound

coming from the direction of cathead no. 4, they saw the second bucket

of concrete was about two-thirds of the way to the top. The static line

went slack. The hoist operator for cathead gantry no. 4 also reported

the loss of tension in the hoist cable at about the same time. He then

applied the brake on the hoist line.

Cathead no. 4 slowly fell toward the inside of the tower togther with the

scaffolding. The collapse of lift 28 commenced and it appeared to progress

toward cathead no. 5. Eventually, lift 28 peeled off with failure radiating

circumferentially in two opposite directions. All scaffolding and forms

fell toward the inside of the tower.

Those who were at the center of the tower took safety under the concrete

truck ramp. All workers who were on the elevated scaffolding were killed

in the collapse.

13

4. LABORATORY INVESTIGATIONS

4.1 INTRODUCTION

The laboratory investigations included testing of concrete specimens under

simulated field conditions to establish the strength and stiffness values of

the concrete of lift 28 at the time of the collapse. The results are also used

in determining the elastic moduli of concrete of other lifts which are needed

for computer analyses of the shell (presented in chapter 6).

Because it was not possible to determine visually the operating condition of

several components of the scaffolding and hoist systems they were examined in

the NBS laboratory. They were also tested to determine their ultimate load

carrying capacities. The components recovered from the Willow Island site

included Williams anchor bolts, two sections of hoist cables, a chain hoist and

a grip-hoist. The results of these component tests are used in evaluating any

component failures which might have triggered the collapse. Examination of possible component failures is presented in chapter 5. Laboratory investigations

on concrete and testing of each of the components are described below.

4.2 CONCRETE TESTS

The purpose of the concrete tests was to establish the compressive, tensile and

bond strengths and the stiffness of the concrete of lift 28 at the time of the

collapse. Although the results of field-cured standard cylinders, made of the same concrete used for lift 28 and tested at 24 and 25 hr , respectively, were known (see table 4.5a and b) , supplementary tests were necessary to establish

the rate of stiffness development and rate of gain in bond strength.

When the collapse occurred, about 18 hours had elapsed since the completion of

concrete placement for lift 28. However, because that part of lift 28 where

cathead no. 4 was located had been completed earlier, it was estimated that the

concrete in that vicinity had been cured about 20 hours prior to the collapse.

Therefore, the 24-and 25-hr test results could not be use directly in the

analysis of the shell.

15
In order to duplicate the concrete used for lift 28 in the laboratory, the

constituent materials for the concrete were obtained from the concrete supplier

who provided the concrete for tower unit no. 2. The materials included cement,

fine and coarse aggregate, fly ash, water-reducing admixture and air-entraining

admixture. These materials will be referred to in the text as the NBS sample.

To assure that these materials, which were obtained six weeks after the collapse

were the same as those used for the lift 28 concrete, the individual materials

were compared with small quantities of the materials obtained by OSHA personnel

from the concrete supplier on May 2, 1978. These materials will be referred to as the OSHA sample. Comparative analyses of the two samples, one obtained by

NBS and the other by OSHA, were carried out by NBS and other laboratories.

4.2.1 Examination of Constituents of the Concrete

Cement

Analysis of both the physical and chemical properties of the NBS and OSHA

samples were made to examine, first, whether these two samples conformed to the

standard requirements given in ASTH C 150 [4.1]2_/ and second, whether the two

samples had the same properties. The ASTM specified values, the test results

of Law Engineering Testing Co .J./ , and the test results of the Cement and Concrete Reference Laboratory, NBS, are given in table 4.1. Except for the 7-day

compressive strength of cement mortar, other physical properties conformed to

the ASTM requirements. It is to be noted that while both samples showed a lower 7-day compressive strength than the ASTM specified value, the NBS sample

showed 16 percent greater strength than the 7-day value of the OSHA sample.

The difference is even greater for the 3-day strength wherein the NBS sample

showed 27 percent greater compressive strength than that of the OSHA sample.

Tne properties of two cement samples were determined by the chemical analysis

method given in ASTM C 114 [4.2]. The analysis was carried out by Law Engineering Testing Laboratory. The results of the two cement sample analyses are given in table 4.2 together with the ASTM standard requirements. It is seen

that although the cement samples were essentially of similar composition, both

NBS and OSHA samples did not conform to the ASTM standard requirements for Type

II cement, in that the silicon dioxide content of the samples was less than the

ASTM minimum value. Therefore, it may be concluded that, based on both physical

and chemical properties, the cement obtained from the concrete supplier would

not have met the ASTM standard requirements for Type II cement. In addition,

while the chemical properties of the two samples compare well, a substantially

greater compressive strength of the NBS sample than that of the OSHA sample

indicates that the strength of concrete made of the NBS sample at early ages

would be greater than that of the OSHA sample cement.

2_/ Numbers in brackets refer to references listed in chapter 10.

J./ See appendix C for the complete reports.

16
Fly Ash

Table 4.3 shows the elemental analyses of fly ash of the NBS and OSHA samples.

The values listed in the table were obtained by means of atomic emission and

absorption spectrometry which was carried out by NBS. The tabulated results

suggest that the two fly ash samples were essentially the same. The total

contents of alumina, iron oxide and silica in each of the samples were calculated and compared in table 4.4 with the requirement given in ASTM C 618 [4.1].

It is seen that both samples conformed to the ASTM standard.

Fine and Coarse Aggregate

Fine and coarse aggregate consisted of natural river sand and gravel. Those

obtained from the concrete supplier were visually compared with the OSHA samples.

No difference between the two samples was noted.

Water-Reducing Admixtures

A water-reducing admixture is a material used for the purpose of reducing the

quantity of water required to produce concrete of a given consistency. These

admixtures increase the slump of concrete for a given water content. Some

admixtures may also retard the setting time of concrete.

The NBS and OSHA samples were compared by means of infrared spectrophotometry

which was carried out by NBS. The chemical composition of the two samples as shown by the analysis indicates that the NBS and OSHA samples were essentially

the same.

An infrared analysis of concrete constituents from a piece of concrete of lift

28 was made to determine the amount of admixture used in the concrete. This was carried out by the Portland Cement Association. The results of the analysis

are given in appendix D. The analysis revealed that the amount of the waterreducing admixture used in the concrete was about the same as the amount specified in the concrete design mix specification, i.e., 3 oz per 100 lb (89 mL per

45.4 kg) of cement.

Air-Entraining Admixtures

Air-entraining admixtures are generally used to improve the durability of

concrete exposed to cycles of freezing and thawing. The workability of fresh

concrete is also improved, and segregation and bleeding are reduced greatly.

The NBS sample of the air-entraining admixture obtained from the concrete supplier was compared with the OSHA sample by means of infrared spectrophotometry.

This analysis was carried out by NBS. The results showed no significant

difference between the NBS and OSM samples.

4.2.2 Determination of Concrete Strengths and Stiffness

After having determined that, except for possibly the cement, the constituents

of the concrete obtained from the concrete supplier by NBS were essentially the same as the ones obtained by the OSHA personnel from the same source shortly

17
after the collapse, a laboratory test program was initiated to determine

various strengths and stiffness values of concrete. Included were tests for

compressive strength, pullout bond strength and modulus of elasticity.

Test specimens were prepared and cured in an environmentally controlled chamber.

Temperature in the chamber was controlled to simulate the temperature conditions at the Willow Island site over the 24 hour period immediately prior to the

collapse. The chamber temperature was controlled using the data obtained from

the Parkersburg airport which is located about 5 miles (8 km) from the Willow

Island site. It should be noted that the airport is situated at an elevation

of about 170 ft above the Ohio River on which the tower was situated. The temperature variation prior to the collapse based on the airport data and the

temperature variation used for curing of concrete specimens are given in

figure 4.1.

The concrete was prepared in several batches using 2 ft3 (0.057 m3) and 10 ft3

(0.28 m3) capacity mixers. The mix proportions were the same as those reported

for the lift 28 concrete (table 4.5). The mix proportions specified for one cubic yard of concrete were scaled down to match the size of mixer. Prior to

mixing, the components of the concrete were prechilled for at least 18 hours

in an environmental chamber at 45°F (7.2°C) to simulate the temperature condition of the components for the lift 28 concrete at the batching plant. To

duplicate the mixing condition, hot water was added for mixing. Table 4.6

gives the data on fresh concrete obtained at the time of specimen preparation.

The following tests were performed:

1) Compressive strength test of 6 x 12 in (150 x 300 mm) cylindrical

specimens

.

2) Bond strength tests using 8 x 8 in (200 x 200 ram) cylindrical pullout

specimens

.

While several series of compressive tests were made to examine the strengthgain characteristics, only one series was carried out for the pullout bond

tests. For the first 24 hour period after casting, all specimens were subjected

to a simulated field temperature condition as described above. Thereafter, the

specimens were cured at 55°F (12.8°C). For 28 day test, a separate set of

three companion compression specimens were cured at 73°F (22.8°C). The actual

temperature of the concrete was recorded periodically by means of a thermocouple

inserted in a 6 x 12 in (150 x 300 mm) cylinder. The specimens cured in the

chamber were tested at 0.5, 1, 2, 3, 5, 7, 14 and 28 days.

Compressive Strength

The compressive tests were performed according to the procedure described in

ASTM C 39 [4.1]. Figure 4.2 shows the compressive test setup. Deformation of

each specimen was measured during the compressive test with a compressometer

such as is described in ASTM C 469 [4.1]. The results of deformation measurements together with strength data were used to determine the modulus of

elasticity.

18
Figure 4.3 shows the results of the compressive strength tests in which the

cotipressive strength is plotted against the maturity of the concrete. The term

"maturity" expressed in units of "degree-day" represents a simple function with

which the combined effect of temperature and time can be related to the gain in

the concrete strength [4.5].

The NBS test data, plotted as squares, are shown in the figure along with the

results of compressive tests carried out by the Ohio Valley Testiag Laboratory

(OVT) and the Pittsburgh Testing Laboratory (PTL) , plotted as triangles, and

the field test data of 6 x 12 in (150 x 300 mm) cylinders for the lift 28 concrete, plotted as circles. The test specimens used by the Ohio Valley Testing

Laboratory and the Pittsburgh Testing Laboratory were made on May 2, 1978, at

the Willow Island site using the concrete delivered by the concrete supplier.

These specimens were field cured for the first 24 hour period and thereafter, in

70°F (21.2°C) lime water. The specimens prepared at the time lift 28 was cast were kept at the base of the tower for the first 24 hour period and subsequently,

moved to a 70°F (21.1°C) fog room.

It is seen in figure 4.3 that the cylinder strengths of lift 28 agree well with

the OVT-PTL data. Comparison with the NBS data shows that at early ages, the

OVT-PTL data fall below the NBS data, while at later ages, the OVT-PTL data lie

above the NBS data. This indicates that the rate of strength gain represented

by the two sets of data are different. The results of the cement analysis discussed in section 4.2.1 suggest that this difference could be attributed to a slower early strength gain exhibited by the cement of the OSHA sample, which

was obtained on the same day as the OVT-PTL specimens were prepared. Because

the compressive strength data of lift 28 matched closely with the strength gain

characteristics of the OVT-PTL data, the OVT-PTL data were used in the analysis

of the shell. However, for the relationship between the compressive strength

and the modulus of elasticity and the pullout bond strength, the NBS test

results were used because they were the only data available.

Modulus of Elasticity

From stress-strain curves of the compressive test the values of the secant

modulus of elasticity are obtained at the stress level of 40 percent of the

maximum compressive stress (0.4 f^) • The stress-strain relationship of con- crete specimens tested at various ages together with the relationship between

the compressive strength and the modulus of elasticity is shown in figure 4.4.

In this figure, the secant moduli are plotted against the square root of the

compressive strength. A linear regression line is shown together with a line

obtained by using the ACI code equation [4.4]. In the analysis of the shell,

different moduli of elasticity determined by the regression line, for each lift were used.

Pullout Bond Strength

The testing procedure for the pullout bond tests was essentially as described

in ASTM C 234 [4.1] except that only the slip of the bar at the free end was measured. Figure 4.5 depicts the test setup. The test specimen consisted of

a 48 in (1.2 ra) long no. 4 (1/2 in, 12.7 mm) deformed reinforcing bar (ASTM

19
615, Grade 60) cast in an 8 x 8 in (200 x 200 ram) waxed cardboard cylinder

mold. Figure 4.6 shows the pullout specimen and mold. The specimens were made

so that the bar was bonded to the concrete for a length of 6 inches (152 mm).

Figure 4.7 shows the pullout test results in which the maximum test load observed

is plotted against the square root of the corresponding compressive strength.

The bond strength of the no. 4 bar obtained according to the regression equation

was used in the strength evaluation of the shell.

4.3 ANCHOR BOLT TESTS

As described in chapter 3, special anchor bolts were used to hold the inside

and outside jumpform beams together, separated by a constant distance, prior to

casting of concrete. After the concrete had set, although friction existed

betv/een the flanges of inside and outside jurapform beams and the shell, the

bolts served as the only positive means of transferring all construction loads

to the shell. The components of the anchor assembly are shown in figure 4.8.

The assembly consisted of two 3/4 in (19 mm) hexhead bolts (ASTM A 494), two

rectangular washers, two tapered threaded couplings and a centerpiece threaded

crimped rod. By adjusting the amount of insertion of the rod into the threaded

couplings, the space between the two jumpform beams was adjusted.

Two anchor assemblies after tension testing to failure are shown in figure 4.9.

In both cases the fracture took place at the middle where the crimp was present.

The average of two anchor assembly tests was 40,350 lb (179.5 kN) . Because it was observed in the field that in many cases the rod was fractured in the thread

coupling through the net cross section at the first thread, a specimen was tested in tension with the anchor assembly embedded in a 12 x 12 x 12 in (305 x

305 X 305 mm) concrete cube. The purpose of this test was to see whether the

failure mode would change if the anchor bolt were tested with concrete encasement. A photograph showing the fractured part is shown in figure 4.10. As can

be seen, the male rod fractured through the net cross section at the first

thread. The maximum test load was 43,200 lb (192.2 kN) . This latter test load

indicated that the strength of an unembedded anchor assembly would give a lower

bound capacity.

4.4 HOIST CABLE TESTS

Two 100 ft (30.4 m) hoist cables were obtained from the Willow Island site.

One section was cut from the hoist end of the cable, and another from the opposite end of the cable wound on the hoist drum. The hoist cable had a diameter

of 9/16 in (14.3 mm) and was made a non-rotating wire rope having two 19 strand

layers

.

For laboratory tests, ten 6 ft (1,8 m) long specimens were prepared, five from

each of the two sections of cable (figure 4,11). Wire rope sockets were attached to the ends of each specimens by means of molten zinc.

Figure 4.12 shows the test setup for the tension test. Of the ten specimens,

eight broke within 2 ft (0.6 m) of one of the socket ends, while the other two

20
broke close to the middle of the specimen. The results of tension tests of the

cables are given in table 4.7.

The tabulated values show that there is no significant difference between the

strength of the cable section which had been subjected to continuous hoisting

stress and that of the cable section which had been worked on the hoist drum,

4.5 CHAIN HOIST TEST

A 2 ton (17.8 kN) capacity chain hoist was used in cathead no. 4 as part of the

counterstatic line. The function of the chain hoist was to adjust the length

of the counterstatic line so as to counterbalance the force in the static cable

and adjust the position of the cathead gantry. Because any slippage in the

chain hoist gear system due to overloading could result in a continuous increase

in the chain length and subsequent dynamic instability of the cathead gantry,

it was desirable that the maximum strength of the chain hoist be determined.

The test setup shown in figure 4.13 duplicates the field condition using the

cable slings and clevises of cathead no. 4. When the test load reached 15,310

lb (68.1 kN) , the internal gear system fractured without the chain slipping.

The fracture of the gear system also shattered the cover plate of the housing

of the chain hoist (figure 4.14).

4.6 GRIP-HOIST TEST

As shown in figure 4.15, the static line was anchored to the ground by means of

a grip-hoist. By jacking the hand lever on the grip-hoist, the static cable

length and tension could be adjusted. Because any slippage of the static line

through the grip-hoist could cause cathead no. 4 to become unstable the griphoist capacity needed to be determined.

Figure 4.16 shows the tension test setup in which the grip-hoist was placed

between the head and platen of the testing machine with the static cable

attached to the lower end. At a maximum load of 19,100 lb (84.96 kN) , the hook

on the lower end broke away from the housing of the grip-hoist (figure 4.17).

Throughout the entire range of test load, no slippage of the static line was observed.

21

5. EXAMINATION OF CONDITIONS OF CONSTRUCTION

ASSEMBLIES AT THE TIME OF THE COLLAPSE

5.1 INTRODUCTION

The construction assemblies which collapsed into the tower may be divided into

three groups: the hoisting system, the scaffolding system and the foriawork sys

tern. Because only 1.0 yd^ (0.8 m-^) of concrete was delivered to the top of the

tower at the time of the collapse, it is reasonable to conclude that no appreci

able forces were being exerted on the formwork system, and that any local failure of the formwork system would not liave been possible. Therefore, this

chapter examines possible failures of those components of the hoisting and

scaffolding systems which could have caused cathead no. 4 to become unstable

under imbalance of forces. Examination of each critical component is supported

by the data collected from site investigations and, where applicable, by test

results

.

5.2 EXAMINATION OF HOISTING SYSTEM

Figure 5.1 identifies the main components of the hoisting system as:

1. Hoist cable

2. Static line

3. Counterstatic line

4. Chain hoist

5. Cathead beam

6. Outside legs

7. Inside legs

8. Outside support brackets

9. Inside support brackets

10. Slide plate assembly

11. Static line anchor - grip-hoist (not shown in figure 5.1, see figure 4.15)

Failure of any of the above items might cause a sudden shift of forces which

would create an unstable condition for the cathead.

23
Hoist Cable

Two of the six hoist cables were broken. These were located at cathead no. 4

and cathead no. 5. A comprehensive examination of the fractured area by the

Federal Bureau of Investigation reported no kinking, crushing or "bird caging"

of the cable and other signs of damage which could have caused distortion in

the cable structure.

As mentioned in section 3.2, both of these cables cut a deep groove in the

upper edge of the remaining shell, lift 27, indicating that they were not broken

until after the debris had started plummeting towards the ground. It is believed

that a substantial amount of weight must have been hanging on the cable to cut

the deep groove and that the break of the cable followed the initiation of the

=ollapse. The laboratory results indicated that the average breaking strength

of the cable of ten specimens is 27,260 lb (121 kN) . Even under an impact factor of 2.0 applied to the hoist load, a maximum operating load which v/ould be

produced by the concrete bucket and the attachments would be about 4,932 lb (22

kN) (table 6.1). Thus, it is reasonable to conclude that this small magnitude

of the hoist load, as compared with the breaking strength, could not have caused

the cables to break, and that the falling weight of the cathead together with

the scaffolding system must have fractured the cable.

Static Line

Examination of all six static cables during the field investigation showed no sign of fracture along their entire length. All of the cables were still

attached to the sliding plate at one end and anchored to the ground through a grip-hoist at the other end.

Counterstatic Line and Chain Hoist

The counterstatic line consisted of a set of two cable slings interconnected in

series by a chain hoist. As presented in chapter 4, the laboratory test

revealed that the weakest link in the counterstatic line was the chain hoist

which had a maximum test capacity of 15,310 lb (68.1 kN) . Because the test

also showed the chain hoist would fail in fracture rather than by slippage in

the gear system, it can be concluded that no slippage occurred due to the hoist

load. This conclusion is also supported by the closeness of the measured length

of the counterstatic line in the field (see section 3.2) with the computed

value based on the reconstruction of the cathead gantry configuration using the

actual member sizes measured in the field.

Cathead Beam

A close examination of the cathead beam of cathead no. 4 revealed that although

the two channels (C8 x 11.5) comprising the beam were bent at the midspan

(figure 3.16), they did not exhibit any signs of buckling between the two cathead legs. This also agrees with a simple elastic analysis of the channel section as a compressive member. The computed buckling load exceeded by a considerable margin the force produced by the hoisting load with an impact factor of

2.0; 23.4 kip capacity vs. 3.58 kip load (104.1 kN vs 15.9 kN)

.

24
Inside and Outside Cathead Legs

Figure 3.15 and 3.16 show the inside and outside cathead legs, respectively.

As discussed in section 3.2, examination of all four individual legs showed no signs of damage or buckling. Support brackets for attachment of the legs to

jacking frames were still intact. No welding failure of the square tubular

legs to the supporting bracket was noted.

Slide Plate Assembly

The slider plate assembly supported the static cable and the inside sheave of

the hoist line. It could slide along a track attached to the inside of the twochannel cathead beam. By turning the large bolt above the beam, the position

of the slide plate could be adjusted. As can be seen in figure 3.13, the slide

plate became detached from the cathead beam. Judging froia the bent shape of

the bolt, the slide plate first separated from the beam with the bolt still

anchored to the clip angle which was attached to the top flanges of tlie cathead

beam. Subsequently, the downward force produced by the falling weight of the

concrete bucket as well as the debris weight on the hoist cable could have bent

the bolt as well as fractured the bolts which connected the clip angle to the

cathead beam. However, the separation of the sliding plate from the cathead

beam is not unique to cathead no, 4. A similar condition was also noted at

catheads no. 1, 5 and 6. Because they were buried under debris it was not possible to ascertain the condition of the slide plate for catheads no. 2 and 3.

Based on this observation, no definite conclusion could be reached as to when

the slide plate for cathead no. 4 separated from the beam.

The possibility of the slide plate being hit by the dead weight balls (figure

5.1), which were attached to the hoist end of the cable, was examined. The

measured length of the hoist cable (see sections 3.2 and 6.2) places the position of the bucket at about 60 ft (18.3 m) below the cathead beam. This agrees

well with eyewitness accounts that the bucket was two thirds of its way to the

top of the tower when the collapse initiated. Therefore, it can be concluded

that the bucket could not have hit the slide plate and caused it to separate

from the cathead beam.

Static Line Anchor

As described in section 3.2, the static line was anchored to the ground through

a grip-hoist. This allowed adjustment of tension in the line. Laboratory

examination of the grip-hoist which anchored the static line for cathead no. 4 revealed that the two brass shear pins in the ratchet mechanism were still in

place. The shear pins limit the amount of torque that could be applied to the

grip-hoist to increase the tension in the static line. It was established from

the manufacturer that these brass pins shear off at a torque level that would

induce a 3500 lb (15.6 kN) tension in the static line. Conversely, if the

tension in the static line is greater than 3500 lb, additional tension could

not be applied to the line by means of torquing the ratchet mechanism. On the

other hand, the gripping mechanism in the grip-hoist had the gripping capacity

in excess of 19000 lb (85 kN) , Since the maximum tension force produced by the

concrete bucket at the critical location is considerably greater than this

25
force {Mil Id vs. 3500 lb) (see section 6.7), it can be concluded that the

tightening of the static line through the ratchet mechanism could not have

pulled the cathead gantry inwards to initiate the collapse.

5.3 EXAMINATION OF SCAFFOLDING SYSTEM

As described in chapter 2, the platform supports were attached to jacking

frames which in turn were supported by jumpform beams. The jumpforra beams were anchored to the shell by means of special anchor bolt assemblies (see section

3.2). Because the flanges of the inside and outside jumpform beams were interconnected by anchor bolts, the anchor bolts in lift 28 could not be pulled out

without either breaking the 3/4 in (19 mm) hexhead bolt, breaking the crimp rod

or tearing the flange of the jumpform beams. Close examination of the jumpform

bearas to which the legs of cathead no. 4 were attached revealed that there were

no signs of excessive deformation at the bolt holes. The minimum specified

capacity of the hexhead bolt according to ASTM A 449 is 40,100 lb (178 kN) and

the average tensile strength of the crimped rod was 40,350 lb (179.5 kN) (see

section 4.3). Since the computed tension force produced in the anchor bolt

assembly by a combination of the loading conditions including the hoist load

with a dynamic amplification factor of 2.0 was only 16,220 lb (71.2 kN) , it was concluded that the bottom anchor bolts in lift 28 could not have failed prior

to the collapse.

5.4 SUMMARY

In the foregoing sections of this chapter the critical components of the

hoisting and scaffolding systems were examined with the support of field laboratory test data. It was shown that each of the components did not fail prior

to the initiation of the collapse. Therefore, it may be concluded that the

collapse did not initiate due to any component failure of the hoisting and

scaffolding systems. In the following chapter, the shell will be analyzed to see whether its capacity was adequate to support superimposed construction

loads

.

26
6. DEFINITION OF LOADS AND SHELL ANALYSIS

6.1 INTRODUCTION AND BACKGROUND INFORMATION

Criteria and rationale for the definition of loads acting at the time of the

collapse are developed in section 6.1 through 6.8 for use in the analytical

investigations of the shell presented in section 6.9. A total of five basic sources of loading (load cases) and the manner of their distribution on the

shell are identified for this analysis as follows:

1. weight of the tower - continuous

2. weight of the scaffolding - 96 ribs

3. live loads on the scaffolds - 96 ribs

4. weight of cathead assemblies - 12 ribs

5. hoisting loads - 2 ribs

Loads attributed to the weight of the tower are generated internally by the

shell analysis programs used in this investigation from the tower geometry and

unit weight of concrete prescribed in the input. The procedure for evaluating

loads from the other sources cited above is discussed in sections 6.4 through

6.7. Assumptions made with regard to dynamic loading in the derivation of

loads are discussed in section 6.8. Geometric considerations for the definition of loads and the mechanism for their conversion into forces directly

applied to the shell are discussed in sections 6.2 and 6.3, respectively.

The various steps used in the conversion of externally applied loads into

equivalent forces acting directly on the shell structure may not be readily

obvious to those having no prior familiarity with the construction scheme used.

Therefore, some background information for the material contained in subsequent

sections is in order.

The scaled line drawing shown in figure 6.1 is a key sketch that should assist

the reader in understanding the construction scheme used at the Willow Island

site. It represents a typical sectional profile of the scaffolding, the cathead assembly, the static and hoist lines, and the concrete bucket. The position of the bucket shown at the top is that assumed at the time concrete is

unloaded into Georgia buggies (not shown) located on the top deck of the inside

27
scaffold. The sketch of the bucket shown by the dotted lines represents an intermediate position during delivery.

The hoist line rides over two sheaves suspended from the cathead beam as indicated. On the outside, it continues on dovra to the ground level (figure

2.17) where it is wound around a hoist drum which controls its movement. The

hoisted loads (concrete bucket, bundled bars, construction hardware, etc.) are attached to the inside terminal of the hoist line, and, in addition, by means

of a pulley mechanism (point K) , ride on the static line GKB supported at points

G and B. The static line provides stability against lateral movement during

delivery of materials. The hoisting loads exert a transverse pull and thereby

induce a tensile force in the static line. As hoisting of materials proceeds

up from the ground level, the tension in the static line increases to a maximum

somewhere close to but below the halfway mark, gradually diminshes thereafter,

and vanishes entirely when the hoist line becomes vertical. At that point the

static line is referred to as becoming slack (i.e., tensionless, or stressfree) but without play.

To determine the hoisting loads acting at the time of the collapse, a brief

note of explanation is needed about the situation just prior to the collapse.

All reinforcing bars and various construction-related equipment were delivered

to the top. The first bucket of concrete was delivered to cathead no. 4.

Worker statements indicated that when the first bucket was delivered to the

top, adjustments in the length of the static line were made at the grip-hoist

below (point I in figure 6.2, section 6.2) to remove excessive slack, if any,

with the assistance of the worker who unloads the bucket at the top. Therefore,

if the first delivery of concrete was used as a trial to adjust (tighten) the

static line for subsequent runs, the force in that line would be higher during

the delivery of the second bucket of concrete to cathead no. 4. The collapse,

in fact, did occur during the second delivery. In the meantime, the first

bucket of concrete was delivered to cathead no. 5 and unloaded into Georgia

buggies. This bucket had not commenced its descent at the time of the collapse,

a situation which was established on the basis of the field data (figure 6.3,

section 6.2). No other hoist lines were in operation at the time of the collapse so that hoisting loads (load case 5) occurred only at the two ribs

flanking cathead no. 4.

In section 6.7, the hoisting loads at cathead no. 4 are calculated based on the

assumption that the static line becomes slack when the bucket reaches its unloading position as shown in figure 6.1, so that its stress-free length is represented

by the solid line GKB. This assumption is further verified by direct measurements of the length of the static line obtained at the site after the collapse,

and used in the subsequent investigation of hoisting loads.

Referring to figure 6.1, it is noted that most of the externally applied loads

(load cases 2 to 5) are transmitted to the pair of opposite jacking frames present at each of the 96 rib locations (the exceptions are discussed in section

6.3). The jacking frames transmit these loads to the jumpform beams which in

turn transmit them to the concrete shell through a series of interconnecting

anchor bolts. The cathead forces, including the weight of the cathead assembly

(load case 4) , and the forces exerted by the static and hoist lines (load case

28
5) , are transmitted to the shell through the cathead leg support brackets

attached to the jacking frames and through the counterstatic line which bears

against a wide flange beam and is attached to the lower diagonal as shown.

Before the analysis of construction loads could be carried out, it was necessary

to gather a substantial amount of information about the construction scheme,

including sizes, lengths and material compositions of the various components

comprising the scaffolding, hoisting and cathead assemblies. A.11 of this information was assembled through numerous site investigations and laboratory testing

of components recovered from the wreckage, supplemented and corroborated by

information from project drawings and worker statements. The source and nature

of this information is cited at the appropriate places in subsequent discussions.

6.2 GEOMETRY

To evaluate the hoisting loads on the tower, it is first necessary to define

the geometric configuration of the hoisting cables used for the delivery of

construction materials to the top of the shell at the cathead locations (figure

2.18). The cables of particular interest are those serving catheads no. 4 and

no. 5 where concrete was being hoisted at the time of the collapse. The required

information was developed on the basis of data obtained from the site following

the collapse.

Part of the field survey data is presented in figure 6.1 through 6.3 Figure

6.1 is a sectional profile (through a vertical plane of revolution) of the

scaffolding and hoisting systems at a typical cathead location. The scaffolding, jacking frames and jumpform beams occur at each of the 96 ribs while the

six catheads, including the static, counterstatic and hoist lines, the skid

board and the wide flange beam, occur at equal intervals (every 16 ribs) along

the circumference of the shell. Each cathead is located between two adjacent

ribs and transmits its loads to these ribs through the counterstatic line

(figure 6.1) and two pairs of inclined legs at the interior and exterior of the

shell, respectively.

In plan, the hoisting layout for catheads no. 4 and no. 5 is shown in figure

6.2. In elevation, the layout for cathead no. 4 (others are similar) is shown

in figure 6.3. These figures are helpful in visualizing the mechanism for the

delivery of materials to the top of the partially-completed tower shell v/here

the casting operations for lift 29 were being carried out. The lifting of materials at cathead no. 4 is controlled by the hoist drum operator stationed at U.

Starting from the drum, the hoist line proceeds toward and around an interior

sheave attached to sheave block T on the ground, toward and around exterior

sheave Q, up and around the two sheaves suspended from the cathead beam and,

before hoisting commences, all the way down to a concrete hopper at B near center A. of the tower. A stationary static line is attached to point G on the

cathead beam at the top. At the lower end, it passes through a clevis secured

to a concrete hopper at B, and terminates at grip hoist I which in turn is

secured to the ground.

The space coordinates of, and distances between, points given in figure 6.2 were established by double triangulation for independent verification of results.

29
Note that the cathead is radially aligned (along AQ) but the horizontal

projection BG (shown dotted) of the chord of the static line is not. Even

though this introduces a horizontal force component of the static line acting

on the cathead beam in the direction perpendicular to the beam axis, its effect

is not significant and has been ignored in subsequent calculations of hoisting

loads (section 6.7). In figure 6.3, dimensions such as those along the cathead

beam, the diameter and pivotal distances of the sheaves, the lengths of the cathead legs, the location of their bases relative to the shell, and the height of

lift 28, were obtained by direct measurement in the field and were verified

against the drawings. The lengths of the hoist lines for catheads no. 4 and

no. 5 shown in figure 6.3 were likewise obtained by measurement in the field.

They represent the length from the hoist drum U to the point of attachment of

the concrete bucket (point K in the figure). The cathead leg dimensions shown

are the projections of the actual lengths of the cathead legs which are inclined

with respect to the plane of the figure. Likewise, line BG in figure 6.3 is

the projection of the chord length of the static line which is inclined relative

to the plane of the figure (i.e., point B lies outside this radial plane).

The results shown in figure 6.3 were obtained from the foregoing data assuming

the cathead beam to be level (see section 6.6 for explanation). They define

the coordinates of the top support points for the static lines for catheads no.

4 and no. 5 as well as their chord lengths and corresponding horizontal and

vertical projections. They also indicate the positions of the concrete buckets

relative to the respective catheads at the time of the collapse, based on the

hoist line lengths measured after the collapse.

6.3 MECHANISM OF LOAD TRANSFER

It will be helpful at this stage to review the mechanism of load transfer to

the shell. Essentially, most items are connected to the jacking frames at the

ribs (figure 6.4). The loads that are transmitted to these frames are indicated

in figure 6.4a and the points at which they act are keyed to the cross-sectional

layout shown in figure 6.4b by circled numbers. Forces F^ through F5 and moment

M are reactions induced by the weight of the scaffolds and live loads acting

upon them. The connection at point 5 is fixed and therefore capable of developing a moment M. As mentioned previously, these forces occur at all of the 96

ribs of the shell. Forces Ap , Aj^, A2 , Bp, and B^ are the cathead leg reac- tion components on the jacking frames (as defined in figure 6.8 and section 6.7)

and forces P^, and are forces induced by the counterstatic line bearing

against the wide flange beam (see also figures 6.1 and 6.4b). These forces as well as force F^ in the counterstatic line (figure 6.8) occur only at the six

pairs of adjacent ribs where the six catheads are located (figure 2.18).

Forces that are not directly transmitted to the jacking frames may be

identified by reference to figure 6.4b. The only force directly applied to the

shell is transmitted through the roller at the lower end of the exterior scaffold. The roller at the lower end of the interior scaffold will tend to sway

away from the shell so that no bearing force can be expected to develop at that

point. On the exterior of the shell, the upper diagonal with the adjustable

ratchet transmits, through connection B]^ at the top, a force which may be

resolved into components normal and parallel to the shell. The normal component

30
is transmitted to the jumpforra beam while the jacking frame recieves the parallel

component through a pinned extension. Similarly, the lower diagonal to which

the counterstatic line is attached exerts a force through its lower connection

F]^. The normal and parallel components of this force are transmitted respectively to the jumpform beam and the jacking frame. In addition, the connection

at this point develops a horizontal pull normal to the plane of the figure due

to the inclination of the counterstatic line with respect to that plane.

The jumpform beams receive forces through two sets of end rollers and a pawl

attached to each jacking frame (see figure 5, appendix A). The pawl is mechanically engaged to one of the lugs on the outstanding flange of the jumpform

beam. The rollers are free to slide along the flange but are constrained

against movement in the normal direction. The forces on the jumpform beam were calculated by treating the jacking frame as a two-span continuous beam on two

exterior roller supports and an interior pin support (at the pawl) as indicated

in figure 6.4a.

Figure 6.5 identifies forces and couples acting on the outstanding flanges of

opposite jumpform beams at a rib. Points A through F (also appearing in figure

6.4a) define the location of the jacking frame supports relative to the jumpform beams. The two exterior diagonals noted above are connected to points B]^ and F]^. Forces F^y Fy, and circumferential moment My are identified by

reference to the local coordinate axes (x, y, z) as shown. Check marks and zeros shown in tabular form indicate respectively those actions (force component

or moment) that can or cannot develop at each point. Actions F^ and My occur in

opposite pairs as noted in the figure, and develop as a result of the inclination

of the cathead legs and counterstatic line relative to the plane of the figure.

The final stage in the conversion of external loads (load cases 2 to 5) into

equivalent actions applied directly to the shell leads to the results tabulated

in figure 6.6. The points of application of these actions, designated by the

letters I, J, K and L (also appearing in figure 6.5), are located on the middle

surface of the shell. Actions occurring at E and F (refer to figure 6.5) are converted into their statically equivalent counterparts and placed at K. Actions

at F^ are converted in a similar manner and placed at L. The resultants of Fy

forces and My moments acting on the cantilever portion of the jumpform beams

(points A, B, C, D and B]^) are distributed equally to the two anchor bolts at I and J. The Fjr forces at I and J induced by the forces acting on the cantilever

portion of the jumpform beams are calculated by treating the anchors at these

junctions as simple supports.

The preceding steps involve certain assumptions regarding the distribution of

forces to the anchor bolts. For instance, the equal distribution of Fy forces

to anchor bolts at I and J implies that the jumpform beams are axially rigid

and no friction can develop at the shell interface. The other extreme situation

would occur if the axial rigidity of the jumpform beams were negligible relative

to that of the shell, in which case, nearly all of the Fy forces from the cantilever portion would be transferred to the top anchor bolt. In reality, the

axial stiffness of the shell is many times that of the jumpform beam and therefore, the actual distribution will be somewhere between the two extremes with

the greater portion of the force going to anchor bolt I. However, the net

31
effect of the two eKtreiae assumptions on the results of the shell analysis is

not significant. In fact, after examining the combined effect of all the

assumptions made in the covnersion of loads from the jurapform beams to the

shell, it was determined that additional reinforcements will not alter the

conclusions of this investigation.

6.4 WEIGHT OF SCAFFOLDING - LOAD CASE 2

The scaffolding assembly, including formwork and steel joists supports (figure

2.5) is axisymmetric so that it transmits equal forces to the 96 peripheral

ribs of the shell (figure 3, appendix A). The procedure for evaluating the

forces at a rib due to the weight of a repetitive scaffolding section (see

figure 6.7) is straightforward and involves no major assumptions. The weights

of the individual components, evaluated from data on sizes and material composition obtained from the site and the drawings, were placed at their respective

centroids and are indicated by vertical arrows in figure 6.7 (arrows with notation are for other load cases). These include planks used for decking and supporting brackets, diagonals and straps, formwork and joists, railings and posts,

safety nets, hydraulic actuators, lines and jacks, stiff backs, jacking frames,

jumpform beams and miscellaneous other hardware. In the analysis, the junctions

of the lower deck brackets and suspended outer straps (figure 6.7) were assumed

pinned so that with the exception of the diagonally braced bracket at the exterior of the shell (second deck from the top), the system was rendered statically

determinate. This bracket, together with the two diagonals was treated as a pinned truss with one degree of redundancy. The final forces on the shell

resulting from case 2 loading are tabulated in figure 6.6.

6.5 LIVE LOADS - LOAD CASE 3

Live loading consisted of the weight of 51 workers, reinforcing bars, Georgia

buggies, electrical and welding equipment, gas cylinders, water drums, tools,

grouting materials, fresh concrete, miscellaneous hoses, wires, cables and

other hardware. The live loads were applied equally to the top decks of the

interior and exterior scaffolds (designated by P-^ in figure 6.7) and were assumed to be uniformly distributed around the periphery of the shell. After

examining all the evidence at hand (field observations at the VJillow Island, W.

Va. and Berwick, Pa. sites, plus worker statements), it was concluded that live

load distributions other than those assumed cannot be reasonably justified so that large concentration of such loading occurring at any one location, including

in the vicinity of catheads no. 4 and no. 5, were ruled out. Furthermore, it

is noted by reference to figure 6.6 that the contribution of live loads to the

total load on the shell is relatively small so that variations in live load

distribution will not significantly affect the shell analysis results. The

procedure for evaluating shell forces induced by live loads (figure 6.6) is the same as load case 2 discussed in the foregoing section.

6.6 WEIGHT OF CATHEAD ASSEMBLY - LOAD CASE 4

The free-body diagram of a cathead in the radial plane is shown in figure 6.8.

Without both the static and counterstatic lines acting (at G and C, respectively), the cathead assembly is rotationally unstable. If, for any reason,

32
the static line were absent, the cathead could still retain its stability

provided the resultant overturning raoraent due to gravity loads about fictitious

point 0 (point of intersection of cathead legs a and b) is counterclockwise

so that the counterstatic line is in tension. Using the values for the gravity

load components W]^ , W2 , W3 and W and their locations specified in figure 6.8,

it can be readily shown that the corresponding overturning raoraent Is, in fact,

counterclockwise if the cathead beam is assumed to be in horizontal alignment.

It can also be shown that it does not take much backward tilting from this

level position (and this can be manipulated by means of the chain hoist - figure 6.7) to cause a rotational collapse of the cathead in the clockwise

direction (outward) in the absence of the static line.

Referring to figure 6.1, it is noted that, with the cathead beam level, the

proximity of the bucket to the top deck of the inner scaffold is sufficient to

permit unloading of the concrete into on-deck Georgia buggies by means of a chute pivoted to the bottom of the bucket (see figure 6.9 for chute detail).

It should be noted that the position of the bucket relative to the deck can be

controlled, in addition to the chain hoist, by the special bolt on the slide

plate assembly to which the static line and the inside sheave of the hoist

line are attached (figure 6.7). The position of this plate was established by

measurement of dimensions of the component parts found among the wreckage at

the site.

The assumptions that the cathead beam was level at the time of the collapse is

based on the foregoing arguments and is further corroborated by the alignments

observed at the Berwick, Pa. site.

The gravity loads which are transmitted to the six pairs of adjacent ribs at

the six cathead locations around the shell are designated in figure 6.7 (shown

in parentheses) and figure 6.8. They consist of the weights of the cathead

legs , W2 , W3 , the weight of the cathead beam assembly W, which includes the

weights of the sheaves, the weight of the hoist line but not the hoisting

loads, the weight of the steel skidboard Pgg, the weight of the wide flange

beam P^^, and tension Tq in the static line due to its own weight. It should

be pointed out that tension Tq depends on the stress-free length of the static

line assumed. The criterion used in determining this length as well as the

corresponding cable forces T, Fp and Fq due to hoisting loads (load case 5) are discussed in section 6.7.

Having established the cathead position and component forces of the assembly,

the corresponding leg reaction components at supports A and B (figure 6.8)

and the forces F^, , P^, and Q^, (figures 6.4 and 6.8) transmitted by the counterstatic line were determined by statics. The final forces on the shell

corresponding to load case 4 are tabulated in figure 6.6.

6.7 HOISTING LOADS - LOAD CASE 5

Once the geometry of the cable profiles are defined (section 6.2), cable

forces and displacements induced by the hoisting loads can be determined. To

33
minimize the human computational effort involved, a computer prograta was developed for that purpose. The program is based on the representation of the static

line by the simple cable model depicted in figure 6.10.

The model assumes the static and hoist lines to be weightless (the weight of

these lines are considered elsewhere - see section 6.6). The hoist line is

very nearly parallel to segment KG of the static line when the bucket is about

midway to the top (location for maximum tension in the static line) and the

program assumes it to be parallel so that force Fq acts in the same direction

as KG for all locations of point K. The input parameters are as defined in

figure 6.10. The fixed input parameters (the values of which are indicated in

the figure) are chord length L, coordinates Xq and Yq of support point G relative to support point B, and the cross-sectional area A and modulus of elasticity E of the static line. The variable input parameters are, position K of the

bucket along the static line as defined by chord length segments Lj^ and L2

,

the normal distance of that point from the chord as defined by parameter D, and

tlie hoisting load W3. Two of the values used for Wg were 2900 lb (1314 kg)

and 5800 lb (2627 kg). They represent, respectively, the weight of the bucket

assembly and concrete (calculations shown in figure 6.9), and twice that

weight to assess the effect of a 100 percent dynamic amplification of hoisting

loads on cable forces.

Figure 6.11 shows in notation form a partial listing of the equations used for

calculating the desired output parameters which are indicated by asterisks.

Not shown are the equations for calculating the cathead leg reaction components

Ap, Bp, A2, B2 and tension F(, in the counterstatic line (figure 6.8), and a refinement introduced in the program to account for changes in the output

parameters resulting from the displacement of support point G due to the elastic

deformation of the cathead assembly.

Note that Sq, the stress-free length o The program was used in an iterative f D to produce the value of Sq that mate

line. The true stress-free length is

f the static line, is an output value,

ashion by adjusting the input parameter

hed the true stress-free length of the

discussed below.

The cathead leg reactions and the force in the counterstatic line are determined

using the free body diagram of the cathead (figure 6.8) under the action of the

static line force T and hoist line forces Fp and Fq . The cathead sheaves are assumed to be frictionless so that Fq = Fp. By reference to figure 6.8, note

that cathead leg reactions A2 and B^ are normal to the plane of the figure and occur in opposite pairs; reactions Ap and Bp act along projections a and b of

the respective inside and outside cathead legs in the plane of the figure; and

reactions k-^ and Bj^ are zero because no intermediate loads (such as, Wj^ , W2 and

W3) are acting in this case (refer to section 6.6). Thus, the resultant

cathead leg reactions are axial. Figures 6.12 and 6.13 should further assist

in visualization of the spatial configuration of the cathead legs, the counterstatic line, and reaction components (also see figure 9, appendix A).

For a given location of point K (figure 6.10), the program outputs two sets

of results. The first set corresponds to the assumption that point G is fixed

in space. The second set is obtained by an iterative process in which

34
successive trials are based on the adjusted position of G corresponding to

the elastic deformation of the cathead assembly under cable forces from the

preceding trial.

Figures 6.13 through 6.15 show the analytical formulations for calculating the

movement of point G attributed to the elastic deformation of the cathead

assembly under the action of the static and hoist line forces. The expression

for the displacement A^, of the counterstatic line under tension F^, is derived

in figure 6.13. In figure 6.14, small rotational displacement geometry is used

to develop the expression for (X^ , Y^) , which represent the movement of

point G, due to the elongation of the counterstatic line. The axial deformations of the cathead legs are smaller than the elongation of the counterstatic

line by about one order of magnitude, and therefore, their contribution to the

movement of point G is negligible. Figure 6.15 shows the formulation for

displacements (Xq, Yq) attributed to the flexural deformation of the cathead

beam. As in the case of the legs, the effect of the axial deformation of the

beam is neglected so that Xq is assumed to be zero. The total displacement

(Xq, Yq) then is obtained by adding the displacements due to the elongation

of the counterstatic line at the top and flexural deformation of the cathead

beam.

In this case, superposition of tension Tq in the static line due to self weight

and tension T caused by loads Wg gives results that are not appreciably different from the tension in the static line (at support G on the cathead beam) that

would be obtained from a solution based on both loads acting simultaneously.

A more refined analysis was performed for a few cases to verify this fact. The

refined analysis considers simultaneously the weight of the cable and the

hoisting loads using an approach similar to that advanced by Harrison [6.1].

The reasons for the small difference between the two analyses are obvious.

The distributed weight of the static line (100 lb or 45.3 kg) is only about

3 percent of the weight of the bucket (2900 lb or 1314 kg). In addition, with

sufficient slack in the static line (this was the case to enable unloading of

concrete at the top of the tower), differences in the results (tension in the

static line as well as tension in the hoist line - figure 6.8) from the two

alternate procedures tend to disappear.

A glance at the results tabulated in figure 6.6 will Indicate that by far the

major portion of the forces acting on the shell is produced by the hoisting

loads (load case 5). In particular, tension T in the static line is the major

source of the bending moments in the shell (hoist line forces Fq and Fq = Fp

remain nearly constant as T varies and balance out so that they contribute

mostly to axial loads Fy but not to normal forces F^ - figures 6.5 and 6.8).

Tension T in turn is quite sensitive to variations in the tautness (or, con- versely, in the amount of slack) in the static line. Thus, the criterion for

estimating the actual stress-free length of the static line at the time of the

collapse would affect significantly the results of the analysis.

As mentioned in section 3.2, the length of the static line of cathead no. 4 (the line was recovered from the wreckage) was measured on two separate occasions in the field. The measured lengths were within 2 in (51 mm) of each

other. The average measured length, after deducting the length of the portion

35
froro grip hoist I to clevis B (figure 6.2), was 219.17 ft (66.85 ra) . With the

backet at the unloading position shovm in figure 6.1 and the static line assumed

tensionless but without play for that position, the stress-free length Sq is

represented by the solid line GKB as shown. This length can be calculated from

the space coordinates of points B and G (figures 6.2 and 6.3) with the known

position of point K (the bucket is hanging vertically down). The computed

length is 219.19 ft (67 .46 m) , which is almost identical to the length measured

in the field.

If a certain amount of play in the static line were allowed when the bucket

is at the unloading position, it would permit workers on the top deck to manipulate the bucket or pull it in, if need be, without encountering resistance

from the static line. However, even without such resistance, a worker will

have to exert a pull of about 400 lb (1780 N) to bring the bucket in a distance

of 1 ft (0.31 ra) . From the layout shown in figure 6.1, with the bucket at

the unloading position, it is apparent that concrete can be discharged without

difficulty into Georgia buggies on the deck by means of the chute projecting

inward above the top rail. Thus, there is no reason to believe that such

manipulation with the bucket was necessary. Furthermore, if there were any

play in the static line at the discharge level, the line would become slack

below that level and the possibility of the bucket swaying and impacting the

scaffold while being pulled up would increase accordingly. Conversely, with

no play at the discharge position, the static line will develop tension at

lower levels and will pull the bucket away from the scaffold thus diminishing

the likelihood of severe impact.

The foregoing arguments led to the conclusion that the field measured length

was the actual length being used at the time of the collapse and, therefore,

was adopted as the basis for the stress-free length used in the calculations

of hoisting loads (load case 5) and the static line tension due to self

weight (load case 4).

The analyis of forces induced by hoisting loads leads to the results tabulated

in table 6.1a. The results obtained by considering the elastic deformation of

the cathead were used to develop the forces for load case 5 tabulated in

figure 6.6. For purposes of comparison only, table 6.1a also shows computer

results obtained by assuming support G to be fixed in space. It should be noted

that the forces in this table correspond to the location of the bucket which

produces maximum tension in the static line. This is about 123 ft (37.5 m)

below the cathead beam which is at variance with the 60 ft (18.3 m) position

estimated from the field data (figure 6.3). From manufacturer's specification

and worker statements the speed at which the hoist line was being pulled in was estimated to be about 10 to 12 ft/sec (3.1 to 3.7 m/s). Hence, it appears

reasonable to assume that failure was initiated when the tension in the static

line was approaching or at, its maximum, and during the next several seconds

while collapse was in progress, it ended up at the 60 ft (18.3 m) mark below

the cathead beam.

Table 6.1b tabulates forces produced by twice the weight of the bucket which is

equivalent to 100 percent impact on the hoist load (or 100 percent dynamic

amplification of the hoist loads). These values are only used to demonstrate

36
that distress conditions would not have occurred in the mechanical system even

if there had been reason to believe that 100 percent impact conditions were present (see chapter 5).

6.8 OTHER EFFECTS

The posture adopted in the foregoing derivations was to use lower bound values

for loads in situations where they were not known precisely. Hoisting loads,

for instance, were treated as though they were stationary because not enough

information was available to assess the inherent dynamic effects of the hoisting

system used. The reasoning behind this approach was that should such loads lead

to structural distress conditions (which turns out to be the case - see section

7), the distinct possibility of higher loads having occurred at the time of the

collapse will not change the final conclusions of this investigation. The

following paragraphs discuss the assumptions made with regard to dynamic

loading

.

Dynamic amplification of loads occurs as a result of construction activities on

the scaffolds (load case 2) and during hoisting of materials (load case 5).

Dynamic amplification of live loads (load case 2) is negligible because, for

the most part, they are stationary and well-dispersed (see section 6.5).

Dynamic amplification of hoisting loads occurs principally as a result of the

initial acceleration of the hoist line to bring the bucket up to speed from

at-rest position on the ground, fluctuations in the speed witli which the bucket

is lifted as the speed of the drum hoist varied, uneven winding of the hoist

line on the drum, or any other changes in speed made by the hoist drum operator

for whatever reasons during the bucket's journey to the top. An increase in

the speed of the bucket will cause a rise in tension in both the static and

hoist lines while a drop in speed will have the opposite effect.

The dynamic loading condition induced by start-up acceleration depends on how

fast the bucket is brought up to speed and the damping characteristics of the

hoisting system. This and the other factors noted above could possibly produce

substantial amplification of hoisting loads. However, not enough information

is available to make a quantitative assessment of these effects with sufficient

accuracy to justify their use as part of the basis for arriving at the

conclusion drawn.

It should be pointed out that the path of the bucket is curvilinear and

therefore, even when moving at a constant speed, the bucket will exert a trans- verse dynamic force (outward normal to the path) on the static line. The path

will be elliptic with foci at supports B and G (figure 6.3) if the cable is

inextensible and nearly so if it is elastic. Placing che bucket at midspan,

assuming a speed of 10 ft/sec or 3.1 m/sec (see section 6.7) and using the

equation of the ellipse yields an incremental tension of about 60 lb (267 N)

in the static ine, v/hich is negligible compared with the 4772 lb (21235 N)

tension (table 6.1) used in this investigation.

37
6.9 ANALYSIS OF SHELL

An analysis of the cooling tower shell subjected to the construction loads

described in the preceding sections was made with SHORE-III [6.2], a finite

element program developed for the static and dynamic analysis of axisymmetric

shells. The cooling tower shell was discretized by a series of curved ring

elements, starting at the top of lift 28. Modeling was consistent with the

shell structure described in chapter 2 of this report. Details of the

SHORE-III analysis and a verification analysis using SAP IV [6.3] are presented

in appendix B. Construction loads were applied to the shell model and the

resultant maximum forces in the shell were determined along with their location.

The finite element model used in the cooling tower analysis is shown along with

an expanded view of the model for the top 3 lifts, in figure 6.16. The tower

structure is discretized by a series of nineteen closed ring elements and one open type element at the bottom which models the column supports. In order to

obtain better stress distributions and properly load the tower, the top two

lifts (28 and 27) were subdivided into eleven elements. The other element

divisions were generally chosen to account for either changes in shell thickness

or changes in the modulus of elasticity of the shell material. Points for load

application were located at the elevations in lifts 28 and 27 where the form

system was bolted to the tower. Figure 6.16 shows the elevations, radii,

thickness, and elastic moduli used for each element in the model. While the

elastic modulus remains constant for an element, the thickness may vary linearly

from the top to the bottom of the element as dictated by the shell geometry.

Material properties (elastic moduli) used in the analysis were obtained from

the test data presented in chapter 4 of this report using the maturity of the

concrete in the various lifts on the day of collapse.

The construction loads presented in figure 6.6 are applied to the shell model

at the load points indicated as distributed line loads by means of a Fourier

Series. This distributed line loading technique is depicted in figure 6.17

for the normal, meridional, and tangential forces, F^ , Fy and F2 , respectively,

and for the meridional moment My. The x, y, z coordinate system shown applies

to the construction loads as described in section 6.3 while the corresponding

u, V, w system is the element coordinate system used in SHORE-III.

The loading function used in the SHORE-III analysis is developed by first

distributing the concentrated loads about the centerline of the jumpform beam

over 0.358 degree circumf erentially (10 in or 254 mm at lift 28) for load cases 2 and 3 and 1.875 degrees circumferentially (4.3 ft or 1.31 m at lift 28)

for load cases 4 and 5. The 10 in (254 mm) distribution width is the surface

contact length between the shell and jumpform beam. This distributed load is

then expanded in a Fourier Series which applies the load at the required points

around the circumference of the tower for the particular construction load case. The larger distribution angle used in load cases 4 and 5 was chosen because it

reduced the number of Fourier series harmonics required to adequately define

the loads. Ideally, a 0.358 degree distribution angle should have been used

for all cases. However, when the 0.358 degree distribution angle was chosen

for load cases 4 and 5, 150 harmonics were required to produce a load function

with an acceptable shape. Increasing the distribution angle to 0.859 degree

38
made it possible to generate a load function with acceptable shape using only

56 harmonics; thereby, reducing computer time and cost by about two thirds.

A comparison analysis using both the 0.358 and 0.859 degree distribution angles

showed that the maximum stresses differed by only a few percent and the stress

distributions were essentially the same. Consequently, the compromise between

distribution angle and computer analysis time is justified. Figure 6.18

illustrates the loading function for a normal force applied to the shell by

the jumpform beams at a cathead. This load would then be repeated at all six

(6) cathead locations in load case 4 to produce a symmetric loading condition.

The development of a convergent Fourier Series with only a few harmonics was found to be a difficult task for load case 5 which is applied only at cathead

no 4. This is because as the number of application points decreases (two

points or ribs for case 5) the number of harmonics required for a convergent

series increases rapidly. In addition, the computer time required for solution

of a load case is related to the number of harmonics in the series. An investigation of the stress distributions that occurred in load case 4 where the six

cathead gantry loads are applied to the shell indicated that the internal

shell forces of interest decayed rapidly to a small value at approximately 20°

from the cathead. The SAP IV verification analysis also exhibited this stress

distribution (see appendix B) . Thus, it was concluded that since the catheads

are 60° apart, the loads applied at a cathead have little influence on the

internal forces at the catheads on either side. Consequently, the loads for

load case 5 were applied at all six cathead locations instead of just at cathead

no . 4

.

The loading functions were developed for each construction load (cases 2 to 5)

and a separate analysis was made for each case, including case 1 which was internally generated by the program. Since SHORE-III is a linear elastic

finite element program, the principal of superposition applies and the results

for each load case may be combined algebraically to find the resultant stresses

for any combination of the construction loads. The effect of using a line load

with a distribution angle rather than a concentrated load is to predict stress

magnitudes which are lower than those that occur in the shell.

Table 6.2 presents the stress resultants in lift 28 at cathead no. 4 for the sum of load cases 1, 2, 3, 4, and 5 at 9=0° and 2°, respectively. Figure 19

shows the sign of conventions for the respective stress resultants. Values at

6=0° occur midway between the jumpform beams to which cathead no. 4 is

attached while the values at 9 = 1.875° occur at the meridian along which the

jumpform beams are attached to the shell.

Figures 6.20 through 6.23 exhibit the distribution of the meridian stress

resultant, N^, the hoop stress resultant, Ng, and the moments, M^,

Mg, with 9 for selected elevations in lift 28 (refer to figure 6.17) for

the sum of load cases 1, 2, 3, 4, and 5. Because they are summetric about the

cathead, the distributions are shown for only one side of the cathead. It

should be noted that the values of the meridional sress at 9 = 1.875°,

figure 6.20 and also table 6.2, increases significantly at elevations of 165.721

and 162.125 ft (50.512 and 49.416 ra respectively) due to the loads being applied

at points I and J (figure 6.6). The shape of the stress distribution for

39
about the circumference of the shell, figure 6.20, can be understood byconsidering the way the loads are applied to the shell. The loads prescribed

in cases 2 and 3 are applied at each rib and the resulting stress distribution

is symmetric about each rib and, furthermore, is identical between all ribs.

On the other hand, the loads prescribed in cases 4 and 3 are applied only at

the cathead locations. The stress distributions illustrated in figure 6.20

reflected the superposition of these cases and the stresses due to load cases 4 and 5 control the shape of the distributions shown. The hoop stress, Nq, as seen from figure 6.21 is highest at the top of lift 28 and decreases with

decreasing elevation. Finally, the maximum values for and Nq occur near

the top bolt in lift 28 at the jumpform beams on either side of cathead 4. A

relative maximum for occurs near the bottom bolt in lift 28 at the jumpform

beams. The maximum moments and Mq also occur at the location of the

jumpform beams (9 = 1.875°) with Mq a maximum near the top of lift 28 and

a maximum near the bottom of lift 28 as shown in figures 6.22 and 6.23.

40
7. INTERPRETATION OF SHELL ANALYSIS

7.1 INTRODUCTION

The overall safety and the reliability of the construction method used at the

Willow Island site depend on the integrity of the partially completed shell

structure to resist all construction loads. This demands that sufficient

strength must be developed in the previously cast shell so that the structural

resistance exceeds the construction load effects by a reasonable margin of

safety. Obviously, the shell did not have adequate strength to resist the

applied load. In order to determine to what extent the applied load effects

exceeded the capacity of the shell, the results of the analysis of the shell

at a number of critical locations are compared with the resistance values

determined by applying existing strength theories. It should be pointed out

that no dynamic amplification of the hoist load was included in the shell

analysis. Dynamic amplification is known to exit in hoisting systems such as the one used at the Willow Island site, and inclusion of the dynamic effects

would have increased the forces in the shell over those determined in this

report

.

7.2 STRENGTH OF THE SHELL UNDER COMBINED AXIAL LOAD AND BENDING MOMENT

For a given cross section and reinforcement, an interaction diagram such as the one shown in figure 7.1 can be constructed in terms of ultimate axial forces as ordinates and ultimate bending moments as abscissa. For various combinations

of axial forces and bending moments, the interaction curve defines a failure

envelope in that all points lying on and outside the interaction curve constitute failure. On the other hand, combinations which fall within the area bounded by the interaction curve represent a safe condition. In this section a number of cross sections in the critical region of the shell are examined using

the interaction diagrams. It was shown previously in the analysis of the shell

that maximum forces in lift 28 would occur in the vicinity of the two ribs

where the legs of cathead no. 4 were positioned. This critical region includes

the section between the two ribs as well as the immediate vicinity along the

ribs

.

41
In developing the Interaction diagrams for the critical cross sections, a shell

section of unit width (1 ft) was treated as a rectangular column. In an actual

shell element the forces and moments occur in adjacent cross sections parallel

and perpendicular to the meridian. By treating a shell element as a column-like

element, the forces and moments in each direction are dealt with separately.

Any effect on the compressive strength of concrete due to biaxial compression

is small [7.2]. Furthermore, the forces acting normal to the reinforcement have

little, if any, effect on the stress that can be developed in the reinforcement.

Thus, the treatment of a shell element as an unaxial element is considered a reasonable and expedient approach.

The cross sectional dimenions of the shell sections used for determining

interaction diagrams are shown in figure 7.2. The vertical section has two

no. 4 bars and the horizontal section has two bars with an equivalent steel area for a 12 in (305 mm) wide cross section. The actual shell cross section

in lift 28 had y/4 vertical brs spaced at 8.7 in (220 mm) on center on each

face and #4 horizontal bars spaced at 12 in (305 mm) on center on each face

(see figure 2.2).

Because the strength-gain characteristics of the lift 28 concrete as shown by

the results of cylinder tests were significantly different from that of the NBS

specimens (figure 4.3), it was decided to use the lift 28 test data to estimate

the concrete strength. The strength of concrete at the time of the collapse,

approximately 20 hours after placement of concrete (section 4.2), is estimated

to be 220 psi (1.52 MPa) based on the results of 24 and 25 hour tests, 283 psi

(1.95 MPa) and 299 psi (2.06 MPa), respectively (table 4.5a and b). The procedure followed to establish the concrete strength is shown in figure 7.3. It is seen that a linear interpolation is made between the time of set (6 hours) and

the time when the field cured cylinders were tested. The time of set was established by the experimental results with cement mortar, with consideration

given to the inclusion of flyash and water reducing admixture in the lift 28

concrete. Approximately the same value of the compressive strength can be

obtained from the strength-maturity plot shown in figure 4.3 with a value of

43 degree F-days being the maturity at the time of collapse. It should be

pointed out that the actual concrete strength at the top of lift 28 could have

been lower than 220 psi (1.52 MPa) due to the temperature condition of lift 28 was different than that of the test cylinders which were cured on the ground

level and due to the rise of free water to the top of the wall both of which

tend to lower the concrete strength. If these factors were taken into consideration, the concrete strength could have been as low as 200 psi (1.38 MPa)

rather than 220 psi (1.52 MPa). However, in the evaluation of cross-sectional

strengths, 220 psi was used.

The reinforcing steel must be embedded adequately in the concrete to develop

its yield strength. The length of embedment required to develop yield is proportional to the square root of the strength of concrete. According to ACI 318-

77 [4.4], at a strength level of 220 psi (1.52 MPa) the embedment length for a no. 4 bar in compression is 40.5 in (1029 mm). Because the ACI code requirement

incorporates a 25 percent increase in embedment length over that required to

develop yield [7.1], the value obtained from the ACI code should be reduced by

42
25 percent to reflect the actual needed embedment lengtVi. Thus, the corrected

compression embedment length for a no . 4 bar becomes 32. A In (823 mm).

The maximum amount of stress that can be developed in the steel was computed

based on available length of embedment. The bars in the circumferential direction, which are in compression, could develop full yield strength of 60 ksi

(414 MPa) except at lap splices. At the locations where lap splices occiar the

maximum stress that can be developed in the bar is limited by the length of

splice. For the splice length of 22 in (559 mm), which was provided for no. 4 bars (figure 2.2), the maximum compression stress in the steel is limited to 60

ksi x 22/32.4 = 40.7 ksi (281 MPa). For the bars in the meridian direction the

available length of embedment varies depending on the distance from the top

edge of lift 28 to cross sections being considered. For a cross section located

32.4 in (823 mm) below the top of lift 28, the steel can develop to yield (60

ksi, 414 MPa). Under combined bending moments and axial forces, some bars were

In tension. For these bars embedment length of 28.8 in (732 mm) was computed

using the pullout test results. However, at all locations where the tension

embedment length would be a governing factor, the magnitude of these moments

and axial forces are relatively low and their combined effects fall well within

the interaction diagram.

For the shell cross sections of unit width shown in figure 7.2 and using the

limiting concrete and steel stresses described above, interaction diagrams for

axial load and bending moment are obtained for critical locations in lift 28

(figure 7.4). In developing the interaction diagram, it was assumed that the

crushing strain of concrete is 0.003 in per i^^/ and that the reinforcing steel

has an elastic-perfectly plastic stress-strain relationship with a nominal

yield strength of 60 ksi (414 MPa).

Figure 7.5 shows an interaction diagram for a cross section resisting meridional

forces (hereafter referred to as the horizontal cross section) for all cross sections 32.4 in (823 ram) below the top of lift 28. Below this point a sufficient embedment length is available for the vertical bars to develop yield. In

lift 28 the meridional force and bending moment are maximum at cross sections

along a line through points BAB as shown in figure 7.4. Maximum force and

moment combinations for cross sections along this line are plotted in figure

7,5 in which the point corresponding to location A is shown as a square and the

point corresponding to location B as a circle. It is to be noted in this figure

that at both locations the cross section is controlled predominately by bending

moment. Because the cathead gantry loads are supported at the rib locations.

!i! In rectangular beam tests, strains 0.003 to 0.004 have been measured near maximum load carrying capacity. Many tests of beams and columns have shown

that a satisfactorily accurate prediction of ultimate strength can be made

using the crushing strain of 0.003 [4.4]. Limited test data are available

for concrete strength of 220 psi (1.52 MPa). However, a more accurate

determination of the concrete strain at maximum resistance was not made in

this investigation because the effect on the interaction diagram is not

large.

43
the bending raoraent is substantially greater at this location than at the center

of the panel, exceeding the capacity of the shell cross section. This indicates

that a compression failure of concrete would initiate at the inside of the

shell at the rib locations and propagate along the circumference at a distance

1 ft (0.3 m) above the bottom of the panel.

The locations of lap splices of the horizontal bars were not known for the cross section resisting hoop forces (hereafter referred to as the vertical cross section). Hence, two interaction diagrams were prepared depending on the maximum

stress that can be developed in the steel. If lap splices of the bars are a sufficient distance away from a section being considered, the steel can develop

yield (60 ksi, 414 MPa) . For cross sections where lap splices occurred, the

maximum stress that can be developed is limited by the length of the splice

(40.7 ksi, 281 llPa) . The interaction diagrams obtained using these two steel

stresses are shown in figure 7.6. These diagrams give the combinations of the

hoop force and bending moment for failure of the cross section. In lift 28,

the hoop force and associated bending moment are maximum at locations C and D.

They are plotted on the interaction diagram for these sections in figure 7.6 in

which the point corresponding to location C are shown in square and the corre- sponding to location D in circle. It is seen in this figure that with point D

lying considerably outside of the interaction curve a compression failure of

concrete at the top of the shell would take place at the rib locations, points

D, due to combined effect of high moment and axial force with or without the

presence of lap splices. At the center of the panel, the shell section is

subjected predominately to axial force.

7.3 STRENGTH OF SHELL UNDER SHEAR

In the presence of axial compression the shear capacity of a reinforced concrete

section increases. This complex interaction between shear and axial compression

is not fully understood. In the absence of a suitable means of determining the

shear strength of a shell section, the expressions in the ACI code [4.4] are used for insight into the strength of the shell in shear. Those expressions

are

:

and not greater than

where

V(> = nominal shear strength provided by concrete

N,, = axial normal load to cross section

44
gross area of section

compressive strength of concrete in psi

b width of section

d distance from extreme compression fiber to centroid of tension

reinforcement

Radial shear values in the vertical and horizontal cross sections of elements

at the top of lift 28 are given in table 7.1 together with corresponding shear

capacities computed in accordance with the above formulas. It may be seen in

the table that the radial shears in the vertical and horizontal sections (Qq

and Q^) are very high at the top of lift 28 along two ribs, while the radial

shears at the center of the panel are either zero or very small. The radial

shear in both the vertical and horizontal sections in the region near the top

of lift 28 along the ribs, points D in figure 7.4, exceed those values computed

by the formulas. Although the calculated shear values may be influenced by the

distribution of concentrated loads along a line, as discussed in section 6.5,

this clearly indicates that relatively high radial shear forces were present

in this region of the shell.

7.4 PROBABLE MODE OF FAILURE

It was shown in the previous sections that maximum stresses occur in the region

of the panel bounded by the tv/o ribs where cathead gantry no. 4 was located.

The analysis showed that a compression failure of concrete would have initiated

in lift 28 along the circumference of the shell at a distance 1.0 ft (0.3 ra) above the bottom of lift 28, the line through points BAB in figure 7.4. In

addition, because both the large hoop forces and moments along the ribs, a band

of the compression failure would also have spread along the rib of the panel.

The presence of high radial shear in the panel could have further weakened the

shell where combination of high axial loads and moments occurred and could have

initiated the failure before the capacity in axial forces and moments was reached

.

Therefore, it may be concluded that the failure of lift 28 resulted from

inadequate strength of the shell section where cathead no. 4 was located to

resist the applied construction loads. The failure was brought about by com- pressive crushing of the concrete due to combinations of axial forces and

moments, and/or cracking due to the high radial shear.

45
<4'-.

v.x.
8. SUMMARY AND CONCLUSIONS

As a result of the cooling tower construction collapse at Willow Island, West

Virginia, the National Bureau of Standards (NBS) was requested by the Occupational Safety and Health Administration (OSHA) to carry out a detailed study

aiming at the determination of the most probable cause of the collapse. In

response to this request, NBS had undertaken a comprehensive field investigation, laboratory tests of construction assembly components and concrete specimens, and chemical analyses of concrete. In addition, mathematical models of

the tower were prepared and analyzed by computer using shell analysis program.

The findings presented in this report are based on the results of these field,

laboratory and analytical studies with other information such as structural

drawings of the tower and OSHA case records.

1. At the time of failure the concrete bucket was in transit from the base of

the tower to cathead no. 4. The measured length of the hoist cable indicated that the bucket was about 60 ft (16.4 m) below the cathead beam.

This also agrees with eyewitness accounts. Therefore, it is believed that

the concrete bucket did not hit the cathead to cause it to fail.

2. Although the hoist cables for catheads no. 4 and 5 were broken, field

observations and the laboratory test indicate that the breakage of the

cables occurred after the onset of collapse of lift 28. Thus, the breakage

of the cables did not trigger the failure.

3. Field and laboratory investigations revealed that the major components of

the hoisting, scaffolding and formwork systems did not fail prior to the

collapse. Thus, the collapse did not initiate due to any component failure

of these systems,

4. Based on the results of field cured cylinder tests, the compressive strength

of concrete of that part of lift 28 where cathead no. 4 was located was estimated to be about 220 psi (1.52 MPa) at the time of the collapse.

5. The analysis of the shell as well as eyewitness accounts indicated that the

collapse initiated in lift 28 at the area where cathead no. 4 was located.

The analysis showed that calculated stress resultants at several points in

47
this area equaled or exceeded the strength of the shell in compression,

bending and shear. Failure at any of these points would have propagated

causing the collapse of lift 28.

These results of the analysis indicate that the most probable cause of the

collapse was due to the imposition of construction loads on the shell befor

the concrete of lift 28 had gained adequate strength to support these loads

48
9. ACKNOWLEDGEMENTS

The authors acknowledge the contributions made by tlie following persons:

Dr. E. 0. Pfrang provided technical guidance and constructive criticism.

Dr. E. V. Leyendecker, Dr. J. R. Harris and Dr. B. Ellingwood critically

reviewed the report.

Mr. R. Williams coordinated the laboratory tests and Messrs. F. Rankin,

J. Owens and T. Ruschell performed the tests.

Mr. C. Mullen coordinated graphic work and prepared figures.

Ms. U. Gray, Mrs. B. Thompson, and Mrs. M. Rarasburg typed the several

drafts of this report manuscript.

Mr. Ray Mele and Mr. John Day carried out the graphic work.

49
.v.;! v;;.:.!;- a/v^s/A

r-:! . .. U:
REFERENCES

1] 1978 Annual Book of ASTM Standards, Part 14, American Society of Testing

and Materials, Philadelphia, PA 19103.

2] 1978 Annual Book of ASTM Standards, Part 13, American Society of Testing

and Materials, Philadelphia, PA 19103.

3] Lew, H. S. and Reichard, T. M. , "Mechanical Properties of Concrete at

Early Ages," Journal of the American Concrete Institute, Vol. 75,

American Concrete Institute, Detroit, Michigan, October 1978.

4] ACI 318-77, "Building Code Requirements for Reinforced Concrete,"

American Concrete Institute, Detroit, Michigan, 1977.

1] Harrison, H. B., "Iterative Nonlinear Structural Analysis," Journal of

the Structural Division, American Society of Civil Engineers, New York,

NY, July 1976.

2] Basu, P. K. and Gould, P. L. , "SHORE-III - Shell of Revolution Finite

Element Program Users' Manual," Res earch Report No. 49, Structural

Division, Washington University, St. Louis, September 1977.

3] Bathe, K. J., Wilson, E. L., and Peterson, F. E., "SAP IV - Structural

Analysis Program for Static and Dynamic Response of Linear Systems,"

Earthquake Engineering Research Center Report No. EERC 73-11, College

of Engineering, University of California, Berkeley, April 1974.

4] Winter, George, Urquhart , L. C, O'Rourke, C. E., and Nilson, /Arthur H. , "Design of Concrete Structures," McGraw-Hill Company, New York, NY,

1973.

1] Orangun, C. 0., Jirsa, J. 0., and Breen, J. E., "A Reevaluation of Test

Data on Development Length and Splices," Journal of the Araerlcaa Concrete

Institute, Vol. 74, American Concrete Institute, Detroit, Michigan,

March 1977.

2] Kupfer, H. , Hilsdorf, H. K. , and Rusch, H. , "Behavior of Concrete Under

Biaxial Stress," Journal of the American Concrete Institute, Vol. 66,

American Concrete Institute, Detroit, Michigan, August 1969.

1] Department of Labor, Bureau of Labor Standards, "Safety and Health

Regulations for Construction," (29 CFR part 1926), Washington, D.C.

51
TABLES AND FIGURES

a

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rH o (U •H O CU bO i-i CO CO O C o CO

CC u c O u c (0 •H s •H rH o J3 CU CU a (1) 4-1 CO CO CO 0) 4-1 rH 4-J CO

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CO •H •H CO CO o 4-1 o •H 0) O c •H < CO 0 CO 1^ H CO hH fU

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52
Table 4.2 Chemical Properties of Type II Cement

Chemical

ASTM

C 150

Wet Chemical Analysis

Elements

%

NBS Sample

%

OSIIA Sample

%

Silicon Dioxide Min. 21.0 20.6 20.0

Aluminum Oxide Max. 6.0 5.3 5.2

Ferric Oxide Max. 6.0 3 .0 3.9

Calcium Oxide — 64.1 63.5

Magnesium Oxide Max. 6.0 2.4 2.1

Sulfur Trioxide Max. 3.0 2.1 2.0

iotax AiKaiies 0.38 0.44

Loss on Ignition Max . j . u 1 .2 1.6

Insoluble Residue iMax . u . / J 0.22 0.25

Tricalciura Silicate 60 60

Dicalcium Silicate 14 12

Tricalcium Alurainate Max. 8.0 7.1 7.1

Tetracalcium Alumiof errite 9 12

53
Table 4.3 Element Aaalysis of Fly Ash

NBS Sample OSHA Sample

Element (in %) (in %)

Aluminum 12.8 11.3

Boron 0.05 0.05

Barium 0. 15 0.20

Carbon 0.50 1.30

Calcium 2.00 3.00

Chromium 0.01 0.01

Copper 0.01 0.01

Iron 9.8 10.5

Potassium 2.00 2.00

Lithium 0.15 0.12

Magnesium 2.00 2.00

Manganese 0.05 0.05

Sodium 0.30 0.30

Nickel 0.01 0.01

Rubidium 0.01 0.01

Silicon 20.9 20.7

Strontium 0.07 0.10

Titanium 0.05 0.50

Table 4,4 Chemical Composition of Fly Ash

NBS Sample OSHA Sample ASTM C 618

Aluminum Oxide 24.1 21.3

Iron Oxide 14.0 14.9

Silicon Dioxide 44.6 44.2

Sum 82.8 80.5 70.0

54
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xTi:

55
Table 4.6 Data on Fresh Concrete

Ran^ Average

1. Unit Weight of Concrete

lb/ft-' (kg/m-';

140.5 - (2251 - 144.4

z J 1 j

;

143.

1

2. Slump

in (ram)

6 - (152 -

9

229)

7.4

(188)

3. Air Content

%

/ c: ^ ,J c; Q q 1

4. Temperature of Concrete

at the Time of Pour

Op (0(.)

66.3 - (17.2 - 66.9

20.6)

66.4

(19.1)

Table 4.7 Results of Tension Tests of Cables

Hoist End Section Drum End Section

1 28.53 kip (126.91 kN) 26.55 kip (118.10 kN)

2 26.82 (119.30) 27.00 (120.01)

3 27.20 (120.99) 26.20 (116.54)

4 27.75 (123.44) 27.45 (122.10)

5 27.40 (121.88) 27.78 (123.57)

Average 27.54 kip (122.50 kN) 27.00 kip (120,01 kN)

56
Table 6.1 Cathead Forces and Reactions Induced by Hoisting Loads - (for load case 5)

(a) (b)

Impact Factor 0% 100%

Wb 2900 5800

lb (kg) (1315) (2630)

So 219.17 219. 17

ft (m) (66.80) (66.80)

Condition of

Support G Fixed Elastic Fixed Elastic

2460.14 2451 .69 4928.35 4931 .73

lb (N) (10943.24) (10905.66) (21922 .31) (21937 .42)

T 5378 .64 4772.26 8860.59 7670.59

lb (N) (23925.37) (21228.06) (39413.85) (34120.47)

Fc 11311 .67 10131 .61 18760.19 16339.25

lb (N) (50316.80) (45067.63) (83449.45) (72680.58)

Ap 18630 .78 16913 .27 31858.23 28461.90

lb (N) (82873.81) (75233.95) (141712.41) (126604.79)

Bp 296.91 496.14 1213.33 1602.59

lb (N) ( 1320 .72) (2206.94) (5397.15) (7128.67)

3453.00 3134.69 5904.57 5275.10

lb (N) (15359.70) (13943.79) (26264.83) (23464.81)

Bz 57.71 96.44 235.84 311.51

lb (N) (256.71) (428.99) (1049.07) (1385.67)

Ll 95.0 100 95.0 100

lb (m) (29.0) (30.5) (29.0) (30.5)

L2 123.56 118.56 123.56 118.56

ft (m) (37.66) (40.41) (37.66) (40.41)

57
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59
Table 7.1 Shear Forces In Lift 28

Loading Cond. Qe Shear

otrengc ii (in pounds)

Qcj) Shear

O t* V*rengL r'\ /T ^ itii (in pounds)

At Center 0 2539 30 2225

At Rib 4388 2594* 8594 2286*

* Capacity of the section is less than shear force.

Q6 = Radial shear in the vertical cross section

Q(}) = Radial shear in the horizontal cross section

1.0 Ibf = 4.44822 N

60
A general view showing completed tower unit no. 1 and unit no. 2

61
62
W CL

#4 @ 8.7" O.C.

4'-7" MIN LAP FOR #8

2'-9" MIN LAP FOR #6

2' 3" MIN LAP FOR #5

1' lO" MIN LAP FOR //4

OUTSIDE r INSIDE FACE

(a) Bar Sizes and Spacing

for Lift 28

(b) Splice Detail for Horizontal Bars

CONSTRUCTION JOINT BETWEEN LIFTS

RIB LOCATIONS

PANEL "C" PANEL "A" PANEL "B" PANEL "C" PANEL "A"

(c) Splice Detail for Vertical Bars

Figure 2.2 Detail of Wall Reinforcement

1 ft = 0.305m

1 in = 25.4mm

63

Cathead gantry

Figure 2.5 A Cross Section through the Formwork and Scaffolding Systems

65
Figure 2.6 Attachment of Jumpfortn Beam to the Exterior of Shell

66

INSIDE JUMPFORM BEAM PAIR

INSIDE JACKING FRAME

Figure 2.8 Attachment Detail of Jumpform Beams and Jacking Frames

68

I Figure 2.10 Forrawork support system

(Note five levels of adjustable joists are used per lift.)

i

70
Figure 2.11 Position of Jurapform Beams prior to Concrete Placement for Lift 27

71
Figure 2.12 Raising of Formwork and Relocation of Jumpform Beam after Casting

of Lift 27

72
Figure 2.13 Position of Jumpform Beams Prior to Concrete Placement for Lift 28

73
Receiving position for

lower jumpform beam

Jacking" of formwork

Elevation 161.125'

Diameter 268.645'

Elevation 156.289'

Diameter 271.142'

Scaffolding system

follows formwork

;ure 2.14 Raising of Formwork and Relocation of Jumpform Beam after Castint

of Lift 28

74
gure 2.15 Position of Jumpform Beams Prior to Concrete Placement

for Lift 29
Figure 2.16 Cathead Gantry Assembly

76
77
Twin drum hoist

Figure 2.18 Location of drum hoist relative to cathead gantries

78
Figure 3.1 Exterior View of Tower Unit No. 2

Hoist cable for cathead no. 1

Figure 3.2 Top of Lift 27 Showing Jagged Edge

79

Figure 3.5 Top of Lift 27 at Cathead Gantry No. 1

Location of cathead

Figure 3.6 Exterior View of Top Portion of the Shell Between Cathead

Gantry No. 1 and No. 2

81
Figure 3.7 Typical Rib in Lift 27

Hexhead bolt

Threaded

coupler

Figure 3.8 Anchor Bolt Assembly

82
ti

' • * —^Localized crushing

Figure 3.9 A Close-up View of the Bottom Bolt in Lift 27 Which

Supported a Cathead Gantry

Figure 3.10 Close-up View of Top of Lift 27 at a Rib Location

83
Figure 3.12 An Overall View of Cathead Gantry

84

Figure 3.15 Inside Legs of Cathead Gantry No. 4
Figure 3.18 Concrete Basket Used for Cathead Gantry No. 4 Which

Was Recovered from Wreckage

87
Wooden cross beams

Figure 3.19 Drum Hoist for Cathead Gantry No. 3 and 4

90

80 -

70 -

60 -

faI 50

E

40

30 -

20

I I I I I I I I I i I I I I I I I I I I I I I I I I I 1 I I I I I I I I r I I I I I I I I I I I

CM

Airport

temperature

30

25

Temperature in curing chamber

' I I I I I I I I 1 1 1 I I I I I I I I I I I I I I I I I I I I 1 I I I I I I I I I 8 10 12 2 4 6 8 10 12 2 4 6 8 10 12 2 4 6 8 10 12 2 4 6 8

26 April p m. 27 April p m. Time (t), days

Figure 4.1 Temperature Variation Used to Cure Test Specimens
Figure 4.2 Compressive test of cylinder

89
90
yfkPa

20

3.0

2.5

COo

2.0 o

UJr 1.5

S 1.0

0.5 -

40 60 80 100 120 140 160

(ACI 318) E=33(145)^^v/f7--_ / V o

E = 0.08685 + 0.05052 yfPT

10 20 30 40 50

Square root of compressive strength in psi

60

180

20

17.5

15

12.5

10

7.5

2.5

70

5000

Stress-Strain Relationship of Concrete

Figure 4.4 Compressive Strength vs. Modulus of Elasticity

91
TEST SPECIMEN

COPPING PLATE

LVDT EXTENSOMETER

HOLDING TRIPOD

No. 4 DEFORMED BAR

SPHERICAL SEAT

TESTING MACHINE

CROSS HEAD

Figure 4.5 Pull-out Test Setup

No. 4 DEFORMED

STEEL BAR

2X4 SUPPORT

SOFT RUBBER TUBING

j

3 In. (7.6cm) 1

6in.(l5.2cm) 1

WIRE CAGE

REINFORCEMENT

48 in Min.

(122 cm)

36 in (9lcm)

8 in. Dia. CARDBOARD

MOLD

lln. l//^y'/Mll>7.>l^/y'/^\

Bin.

(20.3cm)

(2.5cm)

PLASTIC FURNITURE

TIP PLYWOOD BASE

Figure 4.6 Section Through Pull-out Specimen Rack and Mold

92
93
Figure 4.9 Appearance of Williams Anchor Bolts After Tension Test

94
95

Figure 4.13 Test Setup for Chain Hoist

Figure 4,14 Fractured Cover Plate for Gear Box

97
yuve 4.15 Anchorage of static line to ground through grip-hoist

98
Figure 4.16 Test Setup for Grip-Hoist
5.1 Cathead Gantry Assembly

100
Figure 6.1 General Sectional Configuration of Construction Scheme

101
FIELD SURVEY DATA

C: cathead no. 5 static line clevis at concrete hopper

Elevation of C - 105.37 ft (32.13 m)

BD // DE // EF ± AD

Horizontal Distances, ft (m) BD DE AD EF 8.4583 6.5417 1.0625 2.0 (2.5798) (1.9952) (0.3241) (0.61) CE CF CT AF 2.7708 1.5417 92.5625 8.6075

(0.8451) (0.4702) (28.2316) (2.6253)

Coordinates, ft (m)

Location and

Description of Point

Space Coordinates

North South Elevation

A: Center Marker 919.44

(280.43)

1023.39

(312.13)

97.81

(29.83)

B: Cathead 4 Static

Line Clevis at Hopper 1 927.94

(283.02)

1022.66

(311.91)

104.92

(32.00)

T: Sheave Block for Hoist

Lines of Catheads 3 and 4

' 1000.00

(305.00)

1000.00

(305.00)

98.45

(30.03)

N: Cold Joint Top of

| Lift 27 at Cathead 4 1009.23

(307.82)

923.50

(281.67)

264.69

(80.73)

U: Hoist Drum for | Catheads 3 and 4 1079.80

(329.34)

976.84

(297.94)

99.53

(30.36)

Q: Ground Sheave for Hoist

Line of Cathead 4 1040.92

(317.48)

887.84

(270.79)

98.61

(30.08)

V: Sheave Block for Hoist

Lines of Catheads 5 and 6 852.22

(259.93)

973.94

(297.05)

98.49

(30.04)

W: Cold Joint Top of Lift 27 at Cathead 5 869.48

(264.58)

898.75

(274.12)

265.54

(80.99)

R: Ground Sheave for Hoist

Line of Cathead 5 851.94

(259.84)

854.35

(260.58)

98.90

(30.16)

S: Hoist Drum for Catheads 5 and 6 784.79

(239.36)

924.36

(281.93)

99.36

(30.30)

Figure 6.2 Field Survey Data for Location of Hoisting Cables

102
FIELD DATA

LENGTH OF HOIST LINE (UK) AT CATHEAD 4 = 470.4 FT. LENGTH OF HOIST LINE AT CATHEAD 5 = 408 6 FT

COMPUTED RESULTS

N 1001.01 N 874 06

AT CATHEAD 4 AT CATHEAD 5

I ft. = 0 305 m

Figure 6.3 Sectional Elevation at Cathead 4 and Computed Results

of Cable Layout for Catheads 4 and 5

103

\LOC.

FORC^ A B c D E F B, f;

Fx ^/ v/ v/ s/ /

0 0 0 0 0 0

^/ 0 0 \/ 0 v/

My 0 0 / 0 ^/

JUMPFORM

BEAMS

VERTICAL

«- NEAR SIDE JUMPFORM BEAM

(FAR SIDE IS IN OPPOSITE DIRECTION)

1 ft. = 0.305m

Figure 6.5 Forces on the Jumpform Beams

105
y(parallel tc shell)

(normal to shell)

Load Case(2) - Weight of Scaffolding

Loc

Forces\

I J K L

Fx -674 366 91 -226

-2426 -2426 0 0

0 0 0 0

M 0 0 0 0

y

Load Case (4) -Weight of Cathead

-1293 784 -55 232

Fy -928 -928 0 0

213 -123 71 -66

*My 67 66 101 91

Load Case (3) - Live Loads

Loc

I J K L Force s^\

Fx 337 -212 32 0

-970 -970 0 0

z 0 0 0 0

0 0 0 0

Load Case (5) - Hoisting Loads

F^ -19025 12296 426 2964

Fy -1996 -1996 0 0

2151 -2151 707 -1544

*My 798 798 1243 2115

Near side jumpform beam (force on far side beam equal and opposite)

Units: Moment in ft — lb (1.00 ft -s-lb = 1.36 N — m)

Force in lb (1.00 lb = 4.45 N)

Figure 6.6 Summary of Loads Applied to the Shell

106
Cathead beam

Static line Hoist line

Climbing frames

Tubular steel section

Skid board

Steel strap

(Vertical arrows designate forces

of components acting at their

respective centroids —forces in

parentheses occur only at cathead

locations

)

Railing post

Formwork and steel joists

Upper diagonal

Guard rails

Lower diagonal

Planking

Brackets

Rollers

Figure 6.7 Definition of Gravity Loads

107
A, B = supports of inside and outside cathead legs, respectively.

Aj^, Bj^ = cathead leg reaction components in plane of figure and normal to AA and BB^ , respectively.

Ap , Bp = cathead leg reaction components in plane of figure and parallel to AAj^ and BB^, respectively.

A^ , B^ = horizontal cathead leg reaction components of A and B, respectively, normal to plane of figure.

W = resultant weight of cathead beam assembly

T, = tension in static line due to hoisting loads and self weight

Fn, F„ = tension in hoist line (F„ = F ) J ' o Do

F„ = tension in counters tatic line

a = AAj^ = 18 65 ft W3 = 239 lb

b = BBj^ = 17 64 ft

«A

= 77,740°

W = 1071 lb h

= 79.434°

"l = 500 lb = 53.031°

W2 = 449 lb "0

= 12.4°

1 ft = 0,305 m

1 lb = 0,453 kg

Figure 6.8 Free Body Diagram of Cathead under

Gravity and Hoisting Loads

108
WELDED PIPE

CHUTE DETAIL

HOISTING LOADS

VOLUMES:

Bucket

Chute

1/2 yd. Concrete

WEIGHTS:

Bucket

Chute

Concrete

Steel Balls (two)

Misc. Hardware

695 in^ (11.40 x 10^ mm^)

130 in^ (2.13 x 10^ mm^)

13.5 ft^ (0.378 m^)

SCALE:

(.283 X 695)

(.283 X 130)

(150 X 13.5)

TOTAL

Use WB

197 lb (89.2 kg)

37 lb (16.8 kg)

2,025 lb (917.3 kg)

580 lb (263 kg)

65 lb (29 kg)

2,904 lb (1316 kg)

2,900 lb (1314 kg)

<*1

J— 12" (|) BALL

14" ()) BALL

<-0.2 THICKNESS

Figure 6.9 Concrete Bucket Dimensions and Weights

109
CONSTANT INPUT PARAMETERS VARIABLE INPUT PARAMETERS

L = BG = 218.56 ft (66.66 m) L-^ = BQ

A = 0.098 in^ (63.21 mm^) L2 = GQ

E = 13 X 10^ psi (89.7 x 10^ kN/m^) D = KQ

Xq = 115.94 ft (35.36 m) Wg = Bucket Weight

Yq = 185.27 ft (56.51 m)

Figure 6.10 Simple Model for Analysis of Cable

Forces Due to Hoisting Loads

110
^1

=

^2

=

^3 =

^5 =

^1 =

62 =

F =

T =

^1 =

D2 =

s

6 =

F.

-1

tan CD/L-^)

"

tan"-"- (D/L2)

2 - ^

TT — ~ Y

2 2

refer to fig. 6.10

(63 + 64)72

6 - Y2

65 - 61

Wg sin 6-[^/sin ( 3-|^ + 62) = force on static line

Wg sin 62/sin ( 6-|^ + 62) = tension of hoist line

F/(2 cos 6^) = tension in static line

L-j^/sin 63 = length BK, fig. 6.10

L2/sin 64 = length GK, fig. 6.10

^1 ^2 ~ stressed length of static line under tension T

Ts/ (AE + T) = static line elongation

s - 6 = stress-free length of static line

X coordinate of G after elastic deformation of cathead

y coordinate of G after elastic deformation of cathead

force in counterstatic line due to hoisting loads

Cathead leg reaction components in directions

^ shown in figure 6.12 due to hoisting loads

Designates results printed out by computer

Figure 6.11 Procedure for Cable Analysis Program

111
(a) Space configuration

of a cathead leg.

8.79' OUTSIDE LEG (O.L)

(b) View of cathead legs

in tangential plane.

CATHEAD LEGS

COUNTERSTATIC

LINE

1.92"

1 ft. = 0.305 m

(c) Partial shell profile

(d) Configuration of cathead leg

reactions relative to shell.

Figure 6.12 Configuration of Cathead Legs and Reaction

Components Relative to Shell

112
CIq - 68°

PROFILE DEVELOPED LENGTH

FORMULATION OF LOAD-DISPLACEMENT EQUATIONS

Cable: A = 0.154 in^ (99.33 mm^) h

= F^L^/2 (AE)^

E = 13000 ksi (89.7 X 10^ kN/m^) h

= F^L2/2 (AE)2 Chain

A = 0.110 in^ (70.95 mm^) 63 = RL3 sec o/(AE)^

E = 29000 ksl (200 X 10^ kN/m^)

''WF

== F^ (3L^ - 4a)/12EI

(AE)^ = (AE)3 = 2 X 10^ lb (8.9 X 10^ N) 64* = f C-^WF' ^5' ^>

(AE)2 = 3.2 X 10^ lb (14.2 X 10^ N) =

6-^ + 62 '^3 + ^i, R 0.522 F

c = KF^ (K = const.)

1 ft = 0.305

* 64 - displacement of point C due to lateral deflection of WF beam

Figure 6.13 Geometric, Material and Load-displacement

Characteristics of Counterstatic Line
/

/

SMALL DISPLACEMENT GEOMETRY; :

B-|^B2 = bo^''± BB-,^, B-LB3 = A-LA3 = GG-L = = B-|^B2 cos(-I- - 0g)

B2B3 = B-j^B2 sin (-^ - 6g), A2A3 = A-^k^ tan (y - 6^)

Oq = tan"-"- [ (A2A3+B2B3)/A-lB2^] , k^G-^ = k-^G

^1^2 ~ ^3^1 °'G ^2^3 » ^G ~ ^1^3' ^G ~ ^1*^2

^1^2 " '^2'^3 ~

'^l^ °G' ^2^1 ^ tan~-'-(C]^C2/CC2)

6^ = Cj^C sin e^.^^, = KF^ (from figure 6.13)

= (Ae/fic>^G' ^G = (^c/«c)yG**

*

Q'b ~ rotation of BB^^ about point B

** Xq, Yq = displacement of G due to

Figure 6.14 Displacement of Point G due to Axial

Elongation of Counterstatic Line

114
-7^

/

/

/

-7^

2 C 8x 11,5

N. A.

Known

T, Fj5, = Fi3, 6^, = 12.4°

E = 29 X 10^ psl (200 X 10^ kN/m^), I = 65.2 in^ (27.1 x 10^ mm^)

= 5.3 ft (1.6 m), L2 = 3.7 ft (1.1 m)

L3 = 1.8 ft (0.55 m), = 10.5 ft (3.20 m)

Computed

^1 " PilJ/3EI, A2 = (3L-L - P2l|/6EI

M-, = ?-^L-^ , = P2L2, M3 = P3L3

(2n-^ + 2M2 + M3)L/6EI, A3 = L^f

= At + Ao + A-

= 0

*

Xg = + Xg = X^ (see fig. 6.14)

Yg = + Yq ** (see fig. 6.14)

* (Xq, Yq) = displacement of G due to cathead beam deflection

** (XqjYq) = displacement of G due to elastic deformation of cathead

Figure 6.15 Displacement of Point G due to Flexural

Deformation of Cathead Beam

115
EXPANDED VIEW OF TOP 3 LIFTS. F E MODEL

Points

of load

application Element no.

1

Elevation (ft.) Radius (ft)

115171 114337 mat

aim

II2I26

is7:m

1SI2I3

133427 I334II U3ig4

E (ksf) I

Bolt hole location

Element no. Elevation (ft.)

1115 9(9 mm

156213

Ii1444 Mt.eos

141.771

1 1 1 1' I inini ni n 1 1 1 / 1 1 1 1 1 1 1 1 1 1 1 II I nI n1 1 1 / 1 1 in I II 11/

1

Radius (ft.) Tfiickness (ft.) E (ksf) x 10^

133427

134.667

135 9IS 137173 13(437

139.709

117

»>

1.296 \ 967 2.526 \ 667

> < 3.744 \

697 it— 4.632 \

967 >>

4.326 \

II lllll)l II 179.666 II 1 1 II 1 1 1 1 1 1 1 1 1 1 1 1 1 1

1'J 1 1 1 1 1 III! 1 1 rI II 1 1 1 ni 1 1 II I III 1 1 1 1 III fill I

Figure 6.16 Cooling tower finite eleraent iQodel

116
C. Meridional moment

Figure 6.17 Shore-Ill Line Loads

117
( of cathead beam

Figure 6.18 Normal force (F^) at cathead for load case 4

118
119
3

V)O

kiCO

(D

0)

"5

c

o

0)

-7.0

-6.0

(0

.9- -5.0

^ -4.0

-3.0 -

-2.0

£ -1.0

0.0

1.0

LOCATION

OF RIB

/

/

/

/

/

/ 1

(/

i

ELEVATION

165.721 (TOP BOLT IN LIFT 28)

164.337

o 162.125 (BOTTOM BOLT IN LIFT 28)

1

\

1 \

P - A

CENTER OF CATHEAD 4

I 1 1 I

1

-70

-60

-50

-40

g

Z

-30

^

-20

-10

0

10

2 3 4 5

Angle -6, deg

Figure 6.20 Meridional stress resultant at cathead no. 4

120
-35

-30

-25

^ -20

LOCATION

OF RIB

15

10o-o-o

-5

ELEVATION

o 165.969 (TOP OF LIFT 28)

165.721 (TOP BOLT IN LIFT 28)

164.337

O 162.125 BOTTOM BOLT LIFT 28)

--400

-A.

o

CENTER OF CATHEAD 4

-300

E

z

200

•100

1 2 3 4 5

Angle -6, deg

Figure 6.21 Hoop stress resultant at cathead no. 4

121
-90

-80

-70 -

0)

a -ou

IS

c

O -50

c

0) -40

E

o

E

-30

C

0)

E

3

O

o

ELEVATION

o 165.969 (TOP OF LIFT 28)

1 DO .7 21 (lUr dULI in LIrT 2o)

164.337

O 162.125 (BOTTOM BOLT IN LIFT 28

)

if

LOCATION

OF RIB

CENTER OF CATHEAD 4

J L

—

o

I

-20

•15 E

E

-10

-5

0 2 3 4

Angle -G, deg

Figure 6.22 Circumferential moment at cathead no. 4

122
.90

-80

-70

-60

a

3 -50

5-40

c

E -30

o

§-20

6

SI0)

= .10

I

I LOCATION ' OF RIB '

/

/

A'

tLtV A 1 iUN

165.721 (TOP BOLT IN LIFT 28)

A 164.337

0 162.125 (BOTTOM BOLT IN LIFT 28)

\

\

\

\ \

\

\

\ \

\\

\

t \

\

T3

n

10 - ^CENTER OF CATHEAD

4

25

20

•15

E

E

10

-5

1 2 3 4 5

Angle -6, deg

Figure 6.23 Meridional moment at cathead no. 4

123
01 2345678 kN-m

Moment, kip-in

Figure 7.1 An Interaction Diagram under Combined Bending

Moments and Axial Forces

Figure 7.2 Physical Dimensions of Vertical and Horizontal Cross

Sections of a Shell Element Used for Analysis

124
Figure 7.3 Procedure Used to Estimate Compressive Strength of

Concrete at the Time of Failure

CATHEAD GANTRY

No. 4

Figure 7.4 Locations of Maximum Hoop and Meridional Forces

and Bending Moments in Lift 28

125
kN-m

5 6

Center of panel (0°) • Rib (1B75*)

This diagram applies to all sections 32.4 in below the top of lift 28

61.1 kin

-2.7 k, 61.4 k-in

I

B

- 160

140

120

100

- 80

40

40 50

Mi, kip-in

70 80 90

Figure 7.5 Combined Effect of Axial Force and Bending Moment

at Locations A and B

kN-m

50

41.9 k

Center of panel (0°) • Rib(l.875»)

40

34.2 k

30

20

10

No lap splices between two ribs supporting cathead gantry no. 4

Lap splices between two ribs supporting cathead gantry no. 4>

8.1k, 54.6 k-in 'V)

40.0k - in 57.0 k-in

- 120

- 100

80

60

40

4.2k, 64.7 k-in

30 40 M g , kip-in50 60 70 80

Figure 7.6 Combined Effect of Axial Force and Bending Moment

at Locations C and D

126
APPENDICES

Appendix A

Patent of Wall Construction System

United States Patent n-^i Scbeiler

[u] 4,040,774

[45] Aug. 9, 1977

[54] APPARATUS FOR CONSTRUCTING

CONCRETE WALLS

[75] Inventor: Herman Scheller, Eaton. Colo

[73] Assignee: Research-Cottrel, Inc., Bound Brook,

N.J.

[21] Appl. No.: 681,723

[22] Filed: Apr. 29, 1976

[51] Int. Cl.^ E04G 11/04; E04G 1 1/28

[52] U.S. a 425/65; 249/20; 254/139

[58] Field of Search 425/63-65;

249/20-22; 264/33-34; 254/139. 142

[56] References Cited U S PATENT DOCUMENTS

661. .118 11/1900 Nelson 254/13>) 957.521 5/1910 Talbot 249/20

3.472.477 10/1959 Juhl 249/20

3,521.336 7/1970 Rohlf 425/63

3.761.551 9/1973 Ogata el al 264/33

3,779,678 12/1973 Scheller . . 4:5,'b5 Primary Examiner—Francis S Husar

Assisiani Examiner—John McQuadt

Allorney, Agent, or Firm —Harold L Slowell

[57] ABSTRACT

In an apparatus for consiruclina higli-rising. poured

concrete walls, pairs of spaced-apart. upright supports

are preliminarily mounted on a foundation and then detachably attached to both sides of at least a pariially hardened level of concrete wall and al intervals along

the length of the wall for repealed, upward, step-wise

use as the wall is being formed, a plurality of carriages

are mounted on adjacent supports along both sides of

the wall for continuous upward movement as the wall is being cast, and adjustable concrete shaping assemblies

are mounted on the carnages Each assembly opposing

a similar assembly to define a continuous mold into which new concrete is poured on top of previously poured concrete to form the wall

3 Claims, 10 Drawing Figures

A-1
U.S. Patent Aug. 9, 1977 sheet 1 of 6 4,040,774
U.S. Patent Aug. 9, 1977 sheet 2 of 6 4,040,774

FIG 3

A-3
U.S. Patent Aug. 9, 1977 sheet 3 of 6 4,040,774

FIG. 5.

A -4
U.S. Patent Aug. 9, 1977 Sheet 4 of 6 4,040,774
U.S. Patent Aug. 9, 1977 sheet 5 of 6 4,040,774
U.S. Patent Aug. 9, 1977 Sheet 6 of 6 4,040,774

A-7
4,0

1 APPARATUS FOR CONSTRUCTING CONCRETE

WALLS

CROSS-REFERENCE TO RELATED SUBJECT

MATTER

Related subject matter is disclosed and claimed in my

U.S. Pat, No, 3.779,678 granted Dec, 18, 1973. BACKGROUND OF THE INVENTION

Modem concrete wall casting techniques frequently

utilize pairs of spaced-apart, shaping forms held in posi- tion by vanous types of movable suppons, the concrete

being poured between te forms and on top of the section of wall poured earlier and partially set After the last poured concrete has at least partially set, the forms are removed and relocated above the former position and

then the procedure is repeated until the wall is completed SUMMARY OF THE INVENTION

This invention provides a new and improved apparatus for efficient casting of shaped concrete walls and. in the exemplary embodiment, includes a plurality of pairs of spaced-apart, upnght supports extending from the foundation initially and then attached to opposing sides of the wall being formed along the entire length of the

wall, sections of each support being detachable at the bottom of the support and from the wall at intervals and

re-attached at the upper end of the support to form a continuously advancing support as the wall is formed, BRIEF DESCRIPTION OF THE DRAWINGS

FIG 1 is a perspective, stylized view of the base

portion of a cooling tower veil being formed by the

casting apparatus of this invention: FIG, 2 is a vertical section of the apparatus showing

the preferred form of supports and carnage; FIG. 3 IS a top view of the apparatus shown in FIG

2; FIG 4 IS an enlarged, fragmentary side elevation of a form or shaping assembly; FIG, 5 IS an enlarged, vertical section showing the carnage elevating and retaining mechanism, FIG- 6 IS an enlarged, vertical section showing the intermediate form in relation to the carnage; FIG, 7 IS a rear elevation of the sheet portion of the form assernbly; FIG 8 is a fragmentary horizontal sectional view of

the form assembly and related apparatus; FIG 9 is a perspective view of the hoisting apparatus, andFIG. 10 IS a side elevation of the hoisting apparatus DESCRIPTION OF THE PREFERRED

EMBODIMENT

The preferred embodiment of the apparatus is set forth herein, for illustrative purposes, in connection

with the construction of a concrete veil or outside wall of a nature draft cooling tower, which frequently nses 450 feet above the ground and has a diameter of 350 feet

at ground level In such construction, the round concrete wall decreases in diameter as it rises until narrow

section or throat is reached, and then increases in diameter toward the top to form the hyperbolic shape. As

illustrated in FIG 1, the tower veil or wall 10 has an

outer surface 12 and an inner surface 14, and a series of

carriages 18 located entirely around the top of the wall.

M),774 2

During the initial construction of a cooling tower and

after the ground foundation has been laid, pairs of opposed, spaced-apart. generally vertical supports or Hbeams. designated 20 for the inside 14 and 20' for the

5 outside 12, are anchored on the foundation Each sup- pon 20 and 20'. not shown in FIG 1 bun located at each carnage 18, is 20-30 feet long and is composed of

several sections 22 of extruded aluminum, generally H-shaped beams, jointed at their ends by bolted plates

10 24 as shown in FIGS 2 and 5. The beams of each pair are separeated from each other by a distance while will generally be the thickness of the wall 10 being formed, which at the base of a tower is about 30 inches The

beams are spaced apart, as desired, as best shown in

15 FIG 5, by a plurality of horizontal spacers or internally threaded insens 26 attached to the beam flanges 20A by

bolts 26a before concrete is poured between the beams Each pair of beams are separated by about 10 feet from

similar pairs around the wall so that the beams are lo- 20 cated at spaced mtervals completely along both the

inner and outer wall surfaces 12, 14. As the concrete is poured between the beams and the wall increases in height, a lower section of each beam is detached from

the side of the wall 10 by removing bolts 26a from the

25 spacers 26, leaving the spacers in the wall. Then the section is attached to the top of the same beam. As the concrete cures, the lower sections of each support held

tightly against the wall surfaces easily support the car- nages 18 and other structures mounted on them. The

30 beam sections 22 may be lifted into new position by a simple block and tackle or other lifting means mounted

on the carriages 18. The beams 20-20' are also provided

with spaced apart lugs 30 (FIG 2) permanently affixed

to the outer flange 20B of each beam for a purpose later

35 described. A plurality of carriages 18, shown in FIGS. 1, 2, 3, 4, 5 and 6, are separately mounted for vertical movement

on beams 20-20' and adjacent carriages support there between concrete form assemblies 16 which shape the

40 surfaces of the wall being formed. Additionally, as shown on FIG 2, the carnages provide a working plat- form for men and equipment which extends entirely around the wall being formed Each carriage 18, reference being made to FIGS. 2, 4, 45 5 and 6, includes a pair of similar frames 19 composed of

several common structural beams welded and bolted together to form a strong, unitary structure. In detail and as shown in FIG. 2, each frame includes a vertical channel 32, which is welded to a horizontal tubing 46,

50 which in turn is bolted with bolt 132, to channel 34, which IS welded to channel 44. Two adjacent frames 19, one being constructed opposite hand, are jointed to- gether to form each carnage 18 mounted on a beam 20 by bolted connector plates 33 (FIG. 5), angle 38, roller

55 frame 148, and machine bolts 108, 126, 128, 130 and 132, which also serve as axles to other components. A pipe 42 and a thread rod assembly 40 are bolted between

adjacent frames 19, and serve as an adjustable support

to channel 34. Wheels 110 are attached to channel 32

60 and tube 46 of each frame by axles 102, 104, 106 and 108, and ride against opposing sides of the two outer flanges 20B of the beam holding the carnage against the beam. A hydraulic jack 118 (FIG 5) is attached to plate 33 and roller frame 148 in each carriage by bolts and

65 mounts and exerts upward thrust against the carriage when hydraulic pressure is supplied through flexible hose 120. A reciprocating bar 140 is attached to the ram

of the hydraulic jack 118, which in turn has a pivoted

A-8
4,040, 3

pawl 142 which is spring loaded (spring no! shown) lo insure positive engagement between pawl 142 and lug 30 on beam 20-20' Further, wheels 144 are mounted on roller frame 14tl which extends between adjacent beams 32, and are m 5 rolling contact with reciprocating bar 140 to resist lat- eral forces on bar 140 and prevent disengagement be- tween pawl 142 and lugs 30. In addition, a follower pawl 124 welded to a short pipe 125, which in turn is mounted for rotation on a shaft 126 extending between 10 beams 32, engages lugs 30 and is held in engagement by

spring 146. Normally, the carnages progress in an upward direction by extension of pawl 142 until tooth 124 engages a higher lug, 30 followed by withdrawal of pawl 142 for 15 relocation against a higher lug 30, and repeating However, if necessary, a carriage can be lowered by extending hydraulic jack 118 so that the load earned by fol- lower pawl 124 is now transferred to extendable pawl

142 and engaged cleat 30. When the load is thus trans- 20 ferred, pawl 124 can be pivoted so that by slowly re- tracting the jack will allow the carriage to settle and

follower pawl 124 to engage in the next lower cleat 30. A working scaffold is formed by planks 36 resting on beams 34 of adjacent carriages and an outer rail 44 25 between which ropes or planks may be placed Concrete shaping form assemblies 16, generally shown in FIGS 4, 6 and 7 span the distance between

adjacent carriages 18 on both sides of the wall being formed and form and sides of the mold into which fresh 30 concrete is poured Each form assembly 16 includes a forming sheet 64, two end stiffback channels 56, to which lateral, telescoping steel trusses 70 are bolted, with incorporated wedges 74, that strengthen the

trusses 70 when m the proper position Forming sheet 35 64 may be 3/4j plywood for example, or any ridgid material that can be easily cut, or added to, to provide

a smooth casting surface, telescoping trusses 70 provide

lateral support for sheet 64, throughout their effective length, once the incorporated wedges 74 are wedged 40

firmly. The stiffback channels 56 are structural channels that are vertically supported at the bottom by axle 104 which extend beyond beam 32, thru slotted holes 63 (FIGS 6 and 8) to a retainer plate 62. Further, axle 108 45 also extends beyond beam 32 to the retainer plate 62. The form panel 64 is held firmly against the extruded

flange 20A of beams 20—20' This is accomplished by wedging the opposite side of the stiffback channel 56 (FIG 6) with a steel wedge 58 against the extended axle 50 108.Similarly the top of channel 56 forces form panel 64

tightly against the extruded flange 20A of beams 20-20', with a hardwood wedge 60, driven between its opposing side and angle 38. With reference the FIGS 6 and 7, 55 the forming sheet 64 is held vertically by a bracket 66 which merely serves to keep the sheet 64 from falling out when the entire form assembly is unwedged as re- quired while "jacking". As in the case of cooling tower construction, as the 60 tower moves upward in height the diameter decreases

to the throat or neck of the tower, requiring the car- nages 18 lo become closer together. With reference made to FIG. 8 it can be seen that as the carriages 18 move upward and wedges 74 of the steel trusses 70 65 being loosened, they are forced together, or can open

freely as required FIG. 8 also illustrates that while the form assembly 16 has its wedges loosened, the form

774 4

panel 64 is only supported from falling b> bracket 66

and can easily he pulled out be sliding it upward from beiween the support beams 20-20' and the steel trusses 70 for culling. The nature of the extruded aluminum beams 22 that make up the support beams 20-20 permits thai ihcy can

be repeatedly bent or flexed approximately 3 inches over each length without permanent damage

Using this tharactenstii., (FIG 2) a frame 76, comprising two opposite hand wcldmcnts are attached to either half 19 of the carnaec 18 on one side of the wall

10. The frame 76 encases the outer flange 20B of a portion of beams 20-20' w hich is adjacent to firm concrete Similarly a frame 82 is mounted on the lop of

either half 19 of the carnage 18 w hich encases the outer

flange 20B of beams 20-20' above the form enclosure. An arm 78 is bolted to the lower frame 76, and in turn

a stream-bolt ratchet 80 connects to the top frame 82. By extending or contracting the steam-bolt ratchet 80,

lateral pressure is applied sufficient to bend the support beam 20 and properly align the wall 10 being cast. The vertical strength afforded by such a forming

system is also used to best advantage by eliminating excessively high mobile cranes or tower cranes required

in high rise construction, as used in previous forming

systems. Gantries are erected at selected points around

the top of the tower as illustrated in FIGS, 9 and 10. Each gantry is comprised of a beam 88, which is bolted

to four pipe legs 86, which in turn are attached to four separate carriages 18. On the upper side of beam 88, two

sheaves 90 are attached A commercial hoisting engine 93 on the ground is used with these gantries so that the load line 92, originating from the hoisting engine on one

side of wall 10, is supported over the wall on sheaves 90

so that materials can be hoisted on the opposite side of

wall 10 as the arrows indicated in FIG 10. Also, a stationary static line 94, which extends from beam 88 to

a strategic point or concrete hopper on the ground, is used as a guide for the load line 92, which is attached to

a small block 96 so that its sheave wheel rolls on static line 94. The static line 94 serves to make the ground

loading point (not shown) central, which expedites concrete handling from a concrete truck or hopper, as well as keeping the workmen from working beneath the scaffolding along the perimeter of the tower. The two variations in the form surface, or flanges 20A

of jack beams 20 and 20', are shown as in cooling tower

construction, the outer wall surface 12 is generally ndged and the inner wall surface 14 is generally smooth. The plurality of vertical ribs thus formed on

the outside surface 12 of the completed tower serve to induce air lurbulance, thus enhancing heat transfer

In casting the initial courses of the wall, an ordinary

general crane is used to raise the plastic concrete from

the ground to the working scaffold where it is distnb- uted by wheeled carts to the forms around the periph- ery of the structure. It will be appreciated that several cranes may be used simultaneously so that such a large structure can be cast at a reasonable rate. When the structure height exceeds reach of the cranes, a number

of gantries are attached to selected carnages spaced around the wall. These gantries now provide the means

to transport plastic concrete and other materials to the work area. The concrete is raised in buckets (approxi- mately J cu. yd. capacity) by cables running over the

gantries and returning to hoisting engines on the ground. Thus, the necessity for using a large tower

A-9
4,040, 5

crane is avoided and the work can progress much more

quickly and safely than with such a crane

I claim:

1. Apparatus for forming a concrete wall comprising

a. a plurality of generally vertical supports spaced 5 apart along both sides of the wall bemg formed, each support being located opposite to another

support on the other side of the wall and detachably connected to said opposmg support by a plurality of

spacers some of which are contained in the wall 10 previously formed, each support consisting of separate segments detachably joined together at their ends, and the lower portion of each support being

held contiguous the surface of the wall previously formed by said spacers while the upper portion 15 extends above the portion of the wall previously formed.

b. a carriage mounted on each support with means for advancing the carriage upwardly along the supports as the wall is formed, 20

c. a plurality of concrete casting assemblies located along and on both sides of the wall being formed

adjacent its top, each assembly supported at its ends by two carriages on adjacent supports and each

assembly having a generally vertical casting surface 25 extending between the adjacent supports so that the assemblies and the supports together defme a continuous mold extending on top of the wall into which concrete is poured, whereby as the concrete hardens the carriages and assemblies are moved 30 upwardly on the supports to form a new level of

774

6

wall and, at intervals, support segments are de- tached from the bottom of each support and at- tached to the top to provide a continuous track for the carriages,

d. each casting surface compnses the inner surfaces of

generally horizontal, plywood sheets which combine to form a continuous castmg surface,

e. each casting assembly comprising a plurality of

generally horizontal braces positioned contiguous

the outer surface of the plywood sheets to support

the sheets while the concrete is setting against the inner surface of said sheets, and

f each casting assembly compnsing at least a pair of

slotted beams movably earned by adjacent carriages, said plurality of honzontal braces mounted

between said slotted beams which are adapted to be wedged against said adjacent carriages, so that by wedging the slotted beams against the carnages the inner surface of the sheets are drawn into contact

against the adjacent supports thereby providing

said continuous mold.

2. The apparatus defined in claim 1 further including

hoisting means attached to said carnages, a ground

supported hoisting engine, and a hoisting cable connecting said hoisting engine and said hoisting means.

3. The apparatus defined in claim 2 further including

a stationary cable connecting the hoisting means to a central loading point below on which the hoisting cable

is guided.

35

40

45

50

55

60

65

A-10
APPENDIK B

COMPUTER PROGRAMS FOR SHELL ANALYSIS

B.l INTRODUCTION

Failure of the cooling tower shell under the imposed construction loads at the

location of cathead gantry no. 4 was believed to be a plausible cause for the

initiation of the total collapse of lift 28. Thus, an analysis of the shell

structure as it existed at the time of the collapse was made using the con- struction loads. The complexity of the loading conditions as well as the varition in shell thickness and material properties required that the finite element method be employed as the means of analysis. A survey of finite element

programs available for the analysis of this type of shell structure with the

constraints mentioned previously indicated that SHORE-III [6.2] satisfied these

requirements and has been used previously [B.l, B.2] to analyze this type of

structure

.

As a means of verifying the results obtained in the SHORE-III analysis, a second finite element program, SAP IV [6.3] , was selected. SAP IV is a general

finite element program which provides an alternate method for load input and an alternate shell model from those used in the SHORE-III analysis.

This appendix will present a discussion of the pertinent features of the two

finite element programs used in the analysis of the shell and a comparison of

the results obtained from each analysis, for selected loading conditions.

B.2 DISCUSSION OF SHORE-III AND SAP IV

SHORE-III (SHORE) is a finite element program for the linear elastic static and

dynamic analysis of arbitrarily loaded axisymmetric plates and shells. SAP IV

(SAP) on the other hand is a general finite element structural analysis program

for the linear elastic analysis of the three dimensional structural systems.

Although both programs are capable of performing either a static or dynamic

analysis, this discussion will be limited to the static analysis since all

construction loads used in the analysis were treated as being static. Furthermore, since SAP has a rather large element library (truss, beam, plane stress

or strain, three-dimensional solid, pipe, etc.) only the plate/shell element

which was used in the SAP analysis will be discussed.

The capacity of SAP is primarily dependent upon the total number of nodal

points needed to model the shell while SHORE is restricted to a model comprised

of no more than fifty (50) elements. This element restriction for SHORE is not

serious for this analysis since the stress distributions at the lower elevations of the shell do not have a significant effect on the results obtained for

lift 28. Thus, the number of elements used to model the lower portion of the

shell can be reduced.

The shell model is developed for SHORE by discretizing the meridian curve of

the shell with a series of curved rotational ring elements. SAP requires a

B-1
discretization along the meridian and around the circumference of the shell in

order to develop its model. Thus, the SHORE model is composed of a series of

continuous ring elements along the meridian of the shell while the SAP model is

a three-dimensional assemblage of flat plates.

Only a brief description of the elements used in the SHORE and SAP analysis

will be presented. References which present the details on the formulation of

the individual elements are given in the description. The curve rotational ring

element [B.3] used in SHORE has an element stiffness matrix v/hich is derived

from displacement fields that may vary from linear to sixth order and includes

the exact geometry of the shell as well as the effect of transverse shear deformation. The extra coefficients in the higher order displacement fields are eliminated by kinematic condensation at the element level. Sixth order displacement fields were used throughout in this analysis. The element used in

the SAP analysis is a quadrilateral of arbitrary geometry formed from four

compatible triangles. A constant strain triangle [B.4] and a linear curvature

compatible triangle with nine (9) degrees of freedom [B.5] are used to represent the membrane and bending behavior, respectively, of the SAP element. A

central node is located at the average of the coordinates of the four corner

nodes and has six degrees of freedom which are also eliminated by condensation

at the element level.

Each element in both SHORE and SAP have constant material properties (moduli of

elasticity) although the properties may vary as required from element to element in the respective models. The thickness of a particular element in SAP

must be constant, but may change as required throughout the SAP shell model.

The thickness of the SHORE element can vary linearly along the meridian as dictated by the shell geometry.

Various loads including thermal effects may be used as input for both the SHORE

and SAP analysis with distributed loads (gravity, pressure, etc.) and thermal

loads being treated as consistent equivalent nodal loads in both analyses.

However, a major difference exists in the manner in which other loads are input

for analysis. SAP requires that all external loads including moments applied

to the shell structure be input as concentrated loads at nodal points while for

the SHORE analysis all loads are expanded in Fourier harmonics with respect to an element nodal point and the final result is obtained by superimposing the

results of each harmonic.

Finally, both finite element programs solve the resulting set of linear

simultaneous equations for the structural model by a modification of the Gaussian elimination scheme which takes advantage of the symmetric narrow banded

nature of the global matrices used.

B.3 COMPARISON OF RESULTS

Several of the most significant differences between the two finite element

models are considered and comparisons of results are made. The differences

include the effect of element discretization, boundary conditions, the method

of applying the loads, the precision of calculations and the type of elements

used.

B-2
One of the biggest differences between the two analyses lies in the initial

scheme for discretizing the shell. While SHORE uses continuous ring elements

as shown in figure B.l, SAP requires that the rings be broken into numerous

elements. The radial grid adopted is shown in figure B.2. The changes in

radial increment were chosen to reduce the number of nodes required for tlie model in hopes of making the program manageable on the computer while preserving resolution at cathead gantry no. 4 and providing loading points at the

other catheads. As it was, the SAP analysis required raore than eitht hours of

computer time. The effect of having varying radial increments was to introduce

some variations in the solution which could be attributed directly to the

variations in the size of the elements. This can be best illustrated by

considering the gravity loads which should be identical at any angle. Figures

B.3 and B.4 illustrate the radial variation of hoop stress, Nq, and the meridional stress, N(j), respectively, 1.5 ft (.457 m) below the top of lift 28. The

radial variations in the stresses as predicted by SAP are due to the changes in

size of adjacent elements and to the poor aspect ratio of the larger elements.

The smallest elements at cathead no. 4 (elements 126 and 127) are 2,18 ft (0.66

m) long by approximately 0.6 ft (0.18 ra) high for an aspect ratio of 3.63 while

the largest elements near the top (elements 112 and 120) have an aspect ratio

of approximately 58. As can be seen in figure B.3 and B.4 element size change

has more of an effect on N^j, (figure B.3) than the aspect ratio, while the

aspect ratio has more of an influence on Nq (figure B.4). However, figure B.3

and B.4 do illustrate that the stress distribution becomes more uniform near cathead no. 4 where the mesh is finer and the aspect ratio of the elements is

more favorable to obtaining a good solution. In the SHORE analysis, the radial

stress distribution is dependent only upon the equations used to develop the

ring elements. The SHORE and SAP analyses for the gravity load showed good

agreement between N^j, and Nq along the meridian at cathead no. 4. The largest

difference was less than 5 percent.

Different boundary conditions are used in the two models. The SHORE model uses

an open type element developed specifically to represent the columns at the

base of the shell. The effect on the model is to smear the stiffness of the

columns into a ring. In the SAP model the shell at the top of the ring beam

tying the columns together was fixed by prescribing zero displacements. This was done to reduce the number of nodes in the SAP model. Consequently, the

stress distributions within the bottom ring of the shell are quite different

for the two models. However, in element 17 of the SHORE model (see figure B.l)

and the corresponding elements in the SAP model, the stress distributions are similar. The same results are presented in reference B.l where the effect of

different boundary conditions on a hyperboloidal shell were studied. Consequently, the effects of the boundary conditions at the base seem to be far

enough away from the area of interest, the top two lifts, so that none of the

differences in results are due to the differences between boundary conditions

used in the two models.

The method of applying the loads seems to have a great impact upon the results.

SHORE requires that the loads be applied as line loads acting over some finite

length. The length is chosen to allow a reasonably rapid convergence of the

fourier series used to generate the load. SAP requires that the loads be point

loads applied at nodes. For this problem where the loads are essentially

B-3
applied to the shell by bolts, the point load approach is more realistic. The

effect of using the distributed line loads is to cause the stress distribution

to be more uniform near the point of application and the maximum stress

predicted should be lower than the real stress experienced by the shell.

The loading function used in the SHORE analysis is developed by first

distributing the concentrated loads about the centerline of the jumpform

beam. The loads are distributed over 0.358 degrees (10 in or 254 mm at lift

28) for the scaffold loads (load cases 2 and 3) and 0.859 degrees (24 in or 610 mm at lift 28) for the cathead gantry and hoist loads (load cases 4 and 5).

The 10 in (254 mm) distribution width is the surface contact length between the

shell and jumpform beam. This distributed load is then expanded in a Fourier

Series which applies the load at the required points around the circumference

of the tower for that particular construction load. The larger distribution

angle used in load cases 4 and 5 was chosen because it reduces the number of

Fourier Series harmonics required to adequately define the loads. Ideally, a 0.358 degree distribution angle should have been used for all cases. However,

when the distribution angle was reduced from 0.859 to 0.358 degrees, the number

of harmonics required to produce a load function an acceptable shape increased

from 56 to 150 and computer time and costs almost tripled. A sample analysis

was conducted using both the 0.358 and 0.859 degree distribution angles and the

maximum stresses differed by only a few percent. The stress distributions were

also essentially the same. Consequently, it would appear that the compromise

between distribution angle and computer analysis time is justified. Figure B.5

illustrates this distributed line loading for the normal, meridion and tangential forces Fx, Fy , and Fg , respectively and for the meridion moment My,

applied to the shell by the jumpform beams. The bold vectors represent ttie

points loads used in the SAP analysis.

The development of a convergent Fourier Series with only a few harmonics was found to be a difficult task for load cases 4 and 5 which are applied only at

cathead 4. This is because as the number of application points decreases (two

points or ribs for cases 4 and 5), the number of harmonics required for a convergent series increases rapidly. In addition, the required computer time

to analyze a load case using SHORE is related to the number of harmonics. An

investigation of the stress distributions that occurred around the circumference of the shell in load case 4 where the six cathead gantry loads are applied,

for both the SHORE and SAP analysis, indicated that the internal shell forces

decayed rapidly to a small value at approximately 20° from the cathead as shown

in figures B.6 and B.7. Thus, it was concluded that since the catheads are 60°

apart, the loads applied at a cathead have little influence on the internal

forces at the catheads on either side. Consequently, in order to reduce the

number of harmonics necessary to obtain convergence, the loads for load cases 4

and 5 were applied, in the SHORE analysis, at all six cathead locations instead

of just at cathead no. 4. Figure B.8 illustrates the loading function for a normal force applied to the shell by the jumpform beams at a cathead. This

load would then be repeated at all six cathead locations in load case 4 to

produce a symmetric loading condition.

The loading functions were developed for each construction load (cases 2 through 5) and a separate analysis was made for each case, including the

B-4
gravity load which was internally generated by both programs. Since both SHORE

and SAP are linear elastic finite element programs, the principle of superposition applies and the results for each load case may be combined algebraically

to find the resultant stresses for any combination of the construction loads.

Comparison between values of membrane stress, Nq , and bending moment, M , obtained for solutions to load case 4 using SHORE and SAP are illustrated in

figures B.6 and B.7. The comparisons are illustrated for selected elevations

in lift 28 and are typical of results from other load cases. It is interesting

to note that although the loads are distributed over an arc length in the SHORE

analysis, the stress magnitudes predicted compare well with those obtained

using SAP. However, because loads are applied at points in the SAP analysis,

the stresses would be expected to be larger than those obtained using the SHORE

analysis. Several features of the SAP model and solution process may contribute

to the apparent inconsistency. A major feature is that SAP calculates stresses

at the center of the element which essentially represent the average stress in

the whole element. Since even the smallest elements are over two feet wide and

the stress distribution is sharp the peak, stress may be missed by a significant

amount. A second feature involves the precision of the calculations. SHORE

carries out all calculations in double precision on a 32-bit word machine while

the version of SAP used carries out single precision calculations on a 36-bit

word machine. Consequently, roundoff errors may have occurred in the SAP solution process, especially since there were over 6400 equations to be solved and

the band width of the stiffness matrix was 612. Another reason the SAP program

calculates smaller stresses may be the poor aspect ratio of the elements and

the relative size of adjacent elements. Finally, there are differences in the

types of elements used in the two models. SAP uses a plate/ shell element which

provides only for membrane stresses and bending moments. SHORE on the other

hand uses a shell element which accounts for transverse shear and thus provides

for a better estimate of bending moments since the elements are relatively

thick. (0.667 ft, 0.2 m) in lift 28. This appears to be corroborated by the

fact that the comparison between membrane stresses in figure Bo6 is better than

the comparison between bending moments in figure B.7.

B.4 CONCLUSION

Despite the differences in the values of the stresses there are several

encouraging points which arise from the comparison of the two solutions. Both

models gave stress distributions of similar shape for corresponding stress com- ponents. Also, the fact that such different models could lead to the prediction of stresses and moments for which the peak values agreed within a few percent for membrane stresses is encouraging, especially since insufficient time was available to refine the SAP model. Consequently, the SAP model is considered to be a first cut at verification of the SHORE results, while the SHORE

program is designed specifically for the solution of problems involving shells

of revolution.

Based primarily on the difference between the SHORE closed ring element, which

included transverse shear deformation, and the SAP plate/shell approximation

which does not, it is believed that the stress resultant and moments obtained in

the SHORE analysis are a better approximation of the stress levels experienced

B-5
by the shell in lift 28 for the specified construction loads. The actual

stress levels may be higher since the distributed loads used by SHORE to

represent the loads applied to the shell at the jumpforra beam bolts tend to smear the loads over a larger surface area of the shell than actually occurs

and the shell model is not sufficiently able to model the stress distributions

that occur at the bolt locations.

Both the SAP and SHORE models give stress distribution that agree reasonably

well for the type of loads applied in each analysis. Thus, the areas of high

stresses in lift 28 could be predicted from the results of either model.

B.5 REFERENCES

[B.l] Sen, S. K. and Gould, P. L. , " Hyperboloidal Shell on Discrete Supports,"

Journal of the Structural Division, American Society of Civil Engeers,

New York, NY, March 1973.

[B.2] Sen, S. K. and Gould, P. L. , "Free Vibration of Shells of Revolution

Using F.E.M.," Journal of the Engineering Mechanics Division, American

Society of Civil Engineers, New York, NY, Aril 1974.

[B.3] Brombolich, L. J. and Gould, P. L. , "A High-precision Curved Shell Finite

Element," AIAA/ASME 12th Structures, Structural Dynamics and Materials

Conference, American Institute of Aeronautics and Astronautics, American

Society of Mechanical Engineers, Anaheim, CA, April 1971.

[B.4] Felippa, C. A., "Refined Finite Element Analysis of Linnear and

Non-Linear Two-dimensional Structures," SESM Report 66-2, Department of

Civil Engineering, University of California, Berkeley, 1966.

[B.5] Clough, R. W. and Felippa, C. A., "A Refined Quadrilateral Element for

Analysis of Plate Bending," Proceedings 2nd Conference on Matrix Methods

in Structural Mechanics, Wright-Patterson Air Force Base, Ohio, 1968.

B-6
in

o

><

J2

.3 r> cj CO en * 2 S

"5 £2 £2 J2 S2 S7 £

in in

€n ft

LU>

<

X

>

c

'e

mm m ^

m mCM m

^eo

CM <-'e

«D CD tee

9

0}

E

•s 1

03

Figure B.l SHORE-III Finite Element Model for Lifts 26,27 and 28

B-7


GRAVITY LOADS SAP Bf „

Na components 129 128 SHORED! _>

Figure B.4 Comparison of N from gravity load, load case 1

B-10
GRAVITY LOADS SAP IV Nq component

kips/ft

Elevation: 164.351

129 ^2«_127

SHORE III

i3i:!°-i

ft.,. 133^

126

.125

141

142

143-

144

113

-112

152 153 103 — ELEMENT

NUMBERS

Figure B.5 Comparison of Nq from gravity load

B-1

1
C. Meridional moment

Figure B.6 Load functions for SHORE-III and SAP IV programs

B-12
-2.0

-1.8

SHORE m SAP BZ ELEVATION • O 165.544

A 163.116

# o 160.361

- -20

- -15

-25

E—

z

- -10

- -5

10 15 20

Angle -6, deg

25 30

Figure B.7 Comparison of hoop stress resultants for load case 4

B-13
SHORE m SAP Iff ELEVATION • O 165.544

163.116

O 160.361

- -1.75

15 20

Angle -6, deg

25

1.50

-1.25

-1.00

-.75 !

.50

-.25

0.0

25

30

Figure B.8 Comparison of circumferential moment for load case 4

B-14
( of cathead beam

Figure B.9 Loading function for normal force, load case 4

B-15

APPENDIX C

RESULTS OF CHEMICAL ANALYSIS OF CEMENT

LAW ENGINEERING TESTING COMPANY

MAltKIALS irbllNC. ENGINri HS SOIL AND l OUNUAIION I N V ES T K. A I IONS

412 PUiUri Av» , N E • P O Box 13815. Su K ATLANTA, GEORGIA 30324

REPORT OF. CEMENT ANALYSIS

CSiMt: LETCO MC LEAN

P*«>ct: NBS SAf-lPLE #45

LETCO JOB NO. M- 110

Office: Atlanta, Georgia

Date: August 9, 1978

L«b. No. 21 531

B

BRAND OF CEMENT_

PHYSICAL DATA

PLANT SAMPLE RECEIVED 7-15-78

SETTING TIME (Gillmore)

Initial Set:

Final Set:

SOUNDNESS (Expansion)

FINENESS (Surface Area, Blaine)

COMPRESSIVE STRENGTH, PSI

1 Day Break

3 Days Break

7 Days Break

28 Days Break

HOURS

34

0.00

3720

1680

1900

AIR CONTENT (Percent by Volume) 7.2

MINUTES —10

su—

_Sq. Cm./Gm.

CHEMICAL ANALYSIS

Silicon Dioxide (Si02)

Aluminim Oxide (AloOo)

Ferric Oxide (Fe^O^)

Calcium Oxide (CaO;

Magnesium Oxide (MgO)

Sulphur Trioxide (SO3)

Alkalies (Na^O & 0.658 K2O) Loss on Ignition

Insoluble Residue

Trical cium Si 1 i cate (3CaO.Si02)

Dicalcium Silicate (2Ca0.SiO2)

Tricalcium Aluminate (3CaO.AT202)

Tetracalcium Aluminoferrite

(4CaO.Al203.Fe203)

20.0

5.2

1.9 63.5

2.1 2.0

0.44

1.6

"0715 60 IT

12

NOTE: The Type of this cement is not known.

cc: H.S. Lew Washington, D.C. RetpectjuUy nibmitted,

LAW ENGINEERirSG TESTING CO.

C-1
LAW ENGINEERING TESTING COMPANY

MAlFklALS ll bllNC. tNCIfiU I HS SOU AND |-()UNUAIION I N \' LS F I( . A I l( JN S 412 PUttari Av« . N E • P O Box 1381S Sla K ATLANTA, GEORGIA 30324

REPORT OF. CEMENT ANALYSIS

€Si«t: LETCO MC LEAN

Pr^j^.NBS SAMPLE #29

LETCO JOB NO. M- ^0

Office: Atlanta, Georgia

D»te: August 9, 1978

L«b. No. 21 531 -A

BRAND OF CEMENT_?_

PHYSICAL DATA

PLANT SAMPLE RECEIVED7-15-78

SETTING TIME (Gillmore)

Initial Set: Final Set:

SOUNDNESS (Expansion)

FINENESS (Surface Area, Blaine)

COMPRESSIVE STRENGTH, PSI

1 Day Break

3 Days Break

7 Days Break

28 Days Break

HOURS

2

0.00

3710

2140

2200

;iR CONTENT (Percent by Volume) 7.4

MINUTES

4945

_Sq. Cm./Gm.

CHEMICAL ANALYSIS

Silicon Dioxide (Si02)

Aluminim Oxide (AI2O2)

Ferric Oxide (Fe-Oj)

Calcium Oxide (CaO}

Magnesium Oxide (MgO)

Sulphur Trioxide (SO3)

Alkal ies (Na^O & 0.658

Loss on Ignition

Insoluble Residue

Tricalcium Silicate (SCaO.SiOp)

Dicalcium Silicate (2Ca0.SiO27

Tricalcium Aluminate (3CaO.AT20o)

Tetracalcium Aluminoferrite

(4CaO.Al203.Fe203)

K2O)

20.6

5.3 3.0

"647r

2.4

2.1 0.38

1.2

0.22

60

14

7.1

NOTE: The Type of this cement is not known.

cc: H. S. Lew RetpectfuUy $ubmitled,

LAW ENGINEERINq/TES.TING CO.

DAN WELCH

C-2
APPENDIX D

RESULTS OF ANALYSIS OF CONCRETE

Technical Service Report

Project: Slow Setting Concrete Wall Section

Project No.: CT-0477 Date: November 7, 1978 Customer: National Bureau of Standards (Washington, D.C.)

Objective

Determine if Pozzolith 200-N, a water-reducina admixture, was present in concrete placed in a wall section of a thin wall structure at a concentration sufficient to account for an unusually slow setting behavior.

Conclusion

The concrete in question contained a water-reducing admixture,

similar in composition to Pozzolith 200-N, at a level not ex- ceeding that normally recommended by the admixture manufacturer

(3 to 5 fl. oz./lOO lbs. of cement).

Sa:nple Identification and Background Information

A piece of hardened concrete, weighing about 11 lbs. , and three

small bottles of various liquid substances, identified as Pozzolith 200-N, Starch Hydrozylate, and Amine Derivatives Mixture,

respectively, were received from Dr. H. S. Lew, Structures and

Materials Division, Center for Building Technology, on 9/5/78.

Methods of Test

The concrete sample was subjected to chemical analyses to determine the presence and addition level of a water-reducing admixture (ASTM C-494, Type A). The liquid sample identified as Pozzolith 200-N was characterized to obtain its chemical

composition and certain physical properties. The two remaining

liquid samples, identified as Starch Hydrozylate and Amine

Derivatives Mixture, were subjected to infrared analysis to determine the principal ingredients present.

Results and Discussion

The liquid sample identified as Pozzolith 200-N consisted principally of a mixture of corn syrup (Starch Hydrozylate) and

triethanolamine (Amine Derivatives Mixture). The latter com- prised 11% of the admixture formulation as received. Chloride

ion (Cl~) in the amount of 0.24% also was detected. A comparison

of this sample with one analyzed previously by us in October

1975 suggested they were quite similar.

Chemical analysis revealed the concrete sample most li)cely con- tained no more than a normal dose of a water-reducing admixture

similar in composition to Pozzolith 200-N (ASTM C-494, Type A).

In this case, a normal dose is an addition rate of 3 to 5 fluid ounces per 100 lbs. of cement, as recommended by the admixture

manufacturer (Master Builders). A more precise determination of admixture concentration is not possible at this time, unless a calibration curve was to be

prepared employing actual job materials at the levels specified

in the mix design.

L. M. Meyer, Manager

Technical Services Section Chemical Analyses by:

Ig J. R. Polky

CT-0477 Research Chetaist Copy to - A. A. Alonzo iJULCL. J. J. Shideler Associate Research Chemist

D. L. Glochowsl^y

Assistant Research Chemist

D-1
NBS-lUA (REV. 2-ec) U.S. DEPT. OF COMM. 1. PUBLICATION OR 2. Performing Organ. Report No. 3. Publication Date BIBLIOGRAPHIC DATA

REPORT NO.

SHEET (See instructions) NBS BSS 148 September 1982

4. TITLE AND SUBTITLE

Investigation of Construction Failure of Reinforced Concrete Cooling Tower at

Willow Island, WV

5. AUTHOR(S)

H.S. Lew, S.G. Fattal , J.R. Shaver, T.A. Reinhold, and B.J. Hunt

6. PERFORMING ORGANIZATION (If joint or other thon NBS. see in struct/on sj 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE 8. Type of Report & Period Covered

WASHINGTON, D.C. 20234 Final

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City, State. ZIP)

Occupational Safety and Health Administration

U.S. Department of Labor

Washington, DC 20001

10. SUPPLEMENTARY NOTES

Library of Congress Catalog Card Number: 82-600602

Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most si gnificant information. If document includes a si gnificant

bibliography or literature survey, mention it here)

The collapse of the natural-draft hyperbolic concrete cool inq tower unit no. 2 at the

Pleasants Power Station at Willow Island, West Virginia, was investigated by the

National Bureau of Standards. The investigation included on-site inspections, laboratory tests of construction assembly comoonents and concrete specimens, and analytical

studies.

Based on the results of these field, laboratory, and analytical investigations, it was concluded that the most probable cause of the collapse was due to the imoosition of

construction loads on the shell before the concrete of lift 28 had gained adequate

strength to support these loads. The analysis of the shell indicated that the collapse

initiated at the part of the shell in lift 28 where cathead no. 4 was located. It

further showed that calculated stress resultants at several points in that nart equaled

or exceeded the strength of the shell in compression, bending, and shear. The failure

of these points in that part of the shell would have oropagated to cause the collapse

of the entire lift 28.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolon s) collapse; concrete; concrete strength; construction; cooling tower; failure;

hyperbolic shell; shell.

13. AVAILABILITY

^X} Unlimited For Official Distribution. Do Not Release to NTIS ^

Order F rom Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402.

Order From National Technical Information Service (NTIS), Springfield, VA. 22161

14. NO. OF

PRINTED PAGES

156

15. Price

$6.50

USCOMM-DC 6043-P80

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