NATIONAL INSTITUTE OF STANDARDS &

TECHNOLOGY

Research Monnation Center

Gaithersburg, MD 20699

IR 88-3730

RESTRICTED

NBSIR 88-3731

Internal Discontinuity

Sensor Needs for Steel

Report of Workshop

July 16, 1986

Edited by

H. N. G. Wadley and C. D. Rogers

Sponsored by

American Iron and Steel Institute

National Bureau of Standards

1913 1964

Hosted by

U. S. Department of Commerce

National Bureau of Standards

Institute for Materials Science and Engineering

Gaithersburg, MD 20899

June 1988

NBSIR 88-3731

Internal Discontinuity

Sensor Needs for Steel

Report of Workshop

Edited by

H. N. G. Wadley and C D. Rogers

July 16, 1986

Sponsored by

American Iron and Steel Institute

National Bureau of Standards

Hosted by

U. S. Department of Commerce

National Bureau of Standards

Institute for Materials Science and Engineering

Gaithersburg, MD 20899

June 1988

U.S. DEPARTMENT OF COMMERCE, C. William Verity, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

A need has emerged for online detection of internal discontinuities

during continuous casting of steel. Sensing of these defects is generally

considered beyond the sensitivity of existing instrumentation and probably

requires a focused research effort. A workshop was convened on July 16, 1986

at the National Bureau of Standards to assess the status of emerging sensor

technologies and to explore the feasibility of internal discontinuity detection

at various stages during the processing of steel. This report summarizes the

presentations and conclusions of the workshop.

We wish to express our gratitude to J. R. Cook, W. Wilson, M. Linzer

and R. B. Clough for helping to organize and run the workshop and to the

speakers whose presentations provided excellent background and stimulus

for subsequent working group discussions. We would like to also thank

Dr. William E. Dennis of the American Iron and Steel Institute who provided

encouragement and helpful advice for this activity. Finally, we are indebted

to Antonette Nashwinter for assisting in organizing the program and

preparing this report for publication.
.
TABLE OF CONTENTS

1. Introduction 1

2. Formal Presentations 3

Process Control Sensors: Status of AISI Collaborative Programs

H. N. G. Wadley, National Bureau of Standards 5

Recent Progress in the Evaluation of a Pipe/Porosity Sensor

D. Kupperman, Argonne National Laboratory 23

A Pulsed Laser/Electromagnetic Acoustic Transducer Approach to

Ultrasonic Sensor Needs for Steel Processing

G. A. Alers, Magnasonics, Inc. 27

Non-Contact Laser Ultrasonic System for Inspection of Hot Steel

Bodies During Processing

M. Rosen and R. E. Green, Johns Hopkins University 39

Impedance of a Coil in the Vicinity of a Crack

A. H. Kahn, National Bureau of Standards 51

3. Workshop Summaries ___ 59

3.1 Solidification 59

3.2 Intermediate Forming 63

3.3 Finishing 64

SUMMARY AND RECOMMENDATIONS

Operators, engineers and researchers representing American Iron and

Steel Institute member companies met with scientists and engineers from

federal laboratories, universities, and sensor vendor companies to:

• Determine the formation mechanisms of internal discontinuities in

continuous cast steel and their effect on steel quality.

• Identify potential online sensor technologies for internal

discontinuity detection and control strategies for their elimination.

During formal sessions, the Workshop addressed the status of many

emerging sensor technologies with the potential to detect online, internal

discontinuities in steel. During informal workshops, working groups then

convened to discuss the types and origin of discontinuities observed during:

• Solidification

• Intermediate forming

• Finishing

The working groups identified and prioritized the important defects, assessed

the potential of emerging sensor methodologies for their detection and

identified research and development needs.

Steel industry representatives concluded that the top priority problem

was the elimination of surface/ subsurface cracks and inclusions. These

almost always originate in the continuous caster. As cast material undergoes

subsequent rolling/drawing, many of these anomalies result in surface cracks

that are harmful both to quality and productivity. A second priority was

sensing the occurrence of center line segregation /center looseness. The

consensus viewpoint was that each of these defect types could be eliminated

from steel product by adjustment of the continuous caster, provided online

sensors were able to indicate the movement of their formation. The

alternative solution of cropping and discarding unsound material after

offline cold inspection was considered a less desirable expedient because of the

loss of yield and the long time lag before detection of faulty processing

conditions in the caster. While significant strides in statistical process control

and other quality management programs have been made, offline control

scenarios do not allow timely enough detection of internal discontinuities.

The development of online sensors to detect inclusions and cracks

within the caster was determined to be urgently needed for real time quality

v
control. Outputs from such sensors could be used to detect inert gas shroud

failures (the most common source of the most deleterious inclusions) and

cracking problems and facilitate feedback control. Of all the potential sensor methodologies addressed, ultrasonic approaches similar to those developed

for porosity work and also related to those under development for internal

temperature/solidification interface measurement were considered the most

promising. Collaborative programs need to be instituted to prove the

ultrasonic approach and to develop data for the design of a prototype sensor.

vi
1. INTRODUCTION

A Briefing/Workshop on Internal Discontinuity Sensor Needs for

Steel was convened and attended by 50 people on July 16, 1986 at the National

Bureau of Standards, Gaithersburg, Maryland. It was sponsored by the

American Iron and Steel Institute (AISI) and the National Bureau of

Standards (NBS). The aims of the Briefing/Workshop were first to provide

scientists and engineers in industry, universities and government with the

specific needs of sensors for detecting online the formation of internal

discontinuities in steel, and second, to seek immediate input and future

research and development involvement from a broad spectrum of disciplines

to facilitate sensor development. This report is the proceedings of the

presentations and deliberations from the Briefing/Workshop.

The General Research Committee of AISI has identified the

development of Process Control Sensors as an area of research that would

have substantial impact toward improving both productivity and quality in

the steel industry. A January 1981 draft report entitled "Steel Industry

Priorities for Process Control and Sensor Development" identified 537 process

control sensor needs spanning 35 steel processes. A Briefing /Workshop in

July, 1982 addressed in detail four of the sensor needs.

• Automatic detection of pipe and gross porosity in hot steel billets,

blooms or slabs.

• Online inspection of surface defects on hot and cold strip.

• Rapid in-process analysis of molten metal.

• Rapid measurement of temperature distribution within a solid or

solidifying body of hot steel.

Following the workshop, research programs involving cooperation

between AISI member companies, federal laboratories, universities and

vendor companies were initiated for each sensor. These programs are now

reaching the stage of prototype installation (see following summaries in

section 2) and evaluation.

The furthest advanced sensor is that for the automatic detection of pipe

and porosity. An ultrasonic approach has been developed at NBS/Magnaflux

and successfully evaluated at the Argonne National Laboratory (A.N.L.) using

a contact technique. While this contact method of hot ultrasonic inspection

provides an interim (pragmatic) solution to this sensor need, it has been

recognized that ultimately a noncontact system would be preferred.

Noncontact ultrasonic inspection approaches based upon high intensity

lasers /electromagnetic acoustic transduction are beginning to emerge now

1
from the laboratory, and the time may well be right for exploring their

application to the pipe/ porosity sensor need.

Concomitant with this has been the increasing awareness of sensor needs for detecting smaller internal discontinuities such as cracks, inclusions

and center-line segregation which lead to significant reductions in

productivity. Could a modified version of the internal pipe/ porosity sensor

potentially meet these other needs? If so, there would be considerable merit

in developing a coordinated research /development program for a more

general internal discontinuity sensor. The avoidance of duplicated effort, and

the much larger potential market for the generic sensor, would enhance the

cost/risk ratio to the point where a commercial device might emerge from

the sensor vendor industry.

To this objective, a Workshop/Briefing was convened at the National

Bureau of Standards in July 1986. Operators, engineers and research

technologists representing AISI member steel companies met with

researchers from university, government and vendor companies to:

• Identify the most important internal discontinuities and their point

of origination during processing.

• Review candidate sensor methodologies, and suggest promising

research /development directions.

• Encourage the development of a research/development consortium

to pursue development of a sensor.

This report summarizes the proceedings of the meeting.

2
2. FORMAL PRESENTATIONS

3

PROCESS CONTROL SENSORS:

STATUS OF AISI COLLABORATIVE PROGRAMS

H. N. G. WADLEY

Institute for Materials Science and Engineering

National Bureau of Standards

Gaithersburg, MD 20899

INTRODUCTION

The American Iron and Steel Institute (AISI) is presently engaged in several

collaborative research programs whose objective is the development of process

control sensors. These sensors are complex systems based upon the sophisticated

application of advanced technologies featuring high intensity lasers, ultrasound

propagation, high speed digital data acquisition and signal processing. No single

steel industry research laboratory has the resources to tackle these multidisciplinary

research programs. In view of this, these programs feature an unusual level of

cooperation. Cooperation exists between individual AISI member companies

metals industries concerned with other metals such as aluminum, and extends to

several federal laboratories, universities and vendor companies. As these programs

begin to establish varying degrees of feasibility for the proposed sensor technologies,

and the first tangible benefits (prototype sensors) begin to emerge and undergo

industry evaluation, it is opportune to discuss the factors that led to their

conception, development and present status, and to address the future directions

that these programs may take.

HISTORICAL BACKGROUND

The AISI Sensor Programs were conceived in 1979 when the General

Research Committee of AISI identified the development of process control sensors

as an area of research that would substantially impact productivity and quality in the

steel industry. On March 11, 1980, the General Research Committee directed the

establishment of a Task Group to formulate a plan of research and development

relevant to the process control and sensor needs of the steel industry. The purpose

of this group was to develop a framework within which discussion with appropriate

organizations could begin and basic sensor research and development could be

started. The Task Group had very broad industry participation and included

representatives from thirteen AISI member companies.

On January 8, 1981, the Task Group issued a draft report entitled "Steel

Industry Research Priorities for Process Control and Sensor Development" (Ref. 1). In this report, the Task Group critically evaluated sensor applications and identified sensor needs across the entire steel industry. In the analysis, 35 steel processes were

studied, and a total of 537 sensor needs identified. Approximately 70% of the

5
sensors were for applications involving closed- or open-loop control systems. The

majority of the remainder were for periodic monitoring applications, with a small

remaining percentage needed for research techniques.

The Task Group assessed the sensor needs and reached the following

conclusions:

• 40% of the sensor needs could be met adequately with commercially

available equipment.

• 25% of the needs could be partially satisfied with existing equipment.

• 25% of the sensor needs had well established measurement methodologies

but no commercially available equipment.

• 10% of the sensor needs did not have established measurement principles.

The latter two categories were considered potential candidates for

collaborative research programs.

Individual member companies were asked to evaluate these two categories of

needs with respect to present availability, level of interest, and their significance

both for current and proposed future operations. This ranking process reduced the

number of research needs to a list of 18 potential process control sensors that were

then considered, on an industry wide basis, to be the most significant.

From this list of 18 sensors, the AISI assigned top priority to four specific

sensors:

• Rapid In-Process Analysis of Molten Metal.

• Online Inspection for Surface Defects on Hot and Cold Strip.

• Automatic Detection of Pipe and Gross Porosity in Hot Steel, Billets,

Blooms or Slabs.

• Rapid Measurement of Temperature Distribution within a Solid or

Solidifying Body of Hot Steel.

In August 1981, task force units were established for each sensor. These units

were composed of senior scientists and engineers from member companies and

were assigned responsibility for initiating research programs that ultimately would

lead to the development of a sensor and associated analysis methods.

Each task force unit prepared a detailed research prospectus identifying

intended applications for the sensor, benefits of such a sensor to the domestic

6
industry, detailed sensor specifications, and a description of the environment in

which the sensor would be required to operate.

On July 27-28, 1982, a workshop was convened at the National Bureau of

Standards to highlight these four sensor needs (Ref. 2). Presentations by steel

industry research staff were made to a wide range of government, university and

industry researchers and were followed by working group sessions convened to evaluate candidate sensor methodologies. The Workshop concluded with a request

from each of the task force units for research and development proposals from

qualified organizations.

Each task force unit subsequently evaluated all of the proposals received and

developed a research program that, in some cases, involved teaming of proposals.

These research prospectuses were circulated to member steel companies and other

organizations with a shared interest in the research. Individual companies then

grouped together to collaboratively fund the various research programs. Further

cofunding was negotiated with other interested organizations such as the U.S.

Department of Energy (DOE) and the National Bureau of Standards (NBS).

Contracts and cooperative agreements were then signed and the research initiated.

A very important aspect of the research and development to date has been the

development of mechanisms for continued day-to-day involvement of steel

industry research staff with each of the research programs. This has taken many

forms, including frequent program reviews, site visits, and of particular importance

to several of the programs, the extended reassignment of industry research staff to various laboratories to participate in the research itself. This, together with detailed

reporting, has contributed greatly to the success of the technology transfer process.

RAPID INPROCESS ANALYSIS OF MOLTEN METAL

The objective of this program is the development of a sensor for the rapid inprocess chemical analysis of liquid metal and alloys. The sensor is required to

measure only the metal composition and not to sample oxidation products, slag or furnace refractory materials. It must be capable of a rapid multielement readout

(<1 minute), operate without interference to normal liquid metal processing and to be sufficiently robust to withstand the extremely hostile environment of a blast

furnace or refining facility.

The sensor was envisaged to have numerous potential applications in a steel

mill including:

• Continuous iron composition measurement during blast furnace

operation.

• Rapid steel analysis during refining processes.

7
• Sulfur content measurement during iron and steel desulfurization

processing.

• Control of ladle treatments.

• Control of composition at the top of a continuous caster and optimization

of alloy additions during vacuum degassing.

Other unspecified uses /benefits were anticipated for the nonferrous metals

industries, which encouraged development of a generic approach.

The anticipated benefits to the steel industry of these inprocess analysis sensors are very substantial. For example, in 1982, their implementation would

yield $200 million of potential savings from:

• Reduction of heat time. Inprocess chemical analysis would lead to

substantial time savings of 3-7 minutes in heat time during both oxygen

and electric furnace steelmaking thus both improving productivity

(through additional heats in the available time) and reducing operating

costs (by enabling the same number of heats in less time).

• Increase of refractory life. The reduction of heat time would reduce the

length of time that corrosive, end-of-melt slag would be in contact with

refractory linings.

• Reduction of materials costs. Closer composition control during refining

would lead to savings in oxygen, aluminum, flux, and ferro-alloys.

• Reduction of desulfurization costs. Inprocess sulfur analysis during

desulfurization would verify effective removal of sulfur and eliminate

costly and time consuming overtreatment.

The use of inprocess analysis sensors would also reduce the incidence of overblown

heats, and enable better attainment of customer specifications and better quality.

The task force unit evaluated numerous potential approaches and candidate

technologies for an inprocess chemical analysis sensor. These included a variety of

emission spectroscopy techniques coupled with laser ablation /aerosol sampling of

the melt and nuclear activation techniques. The approach considered most

promising was the combination of laser ablation as a means of sampling, coupled

with emission spectroscopy as a method of analysis.

In this approach, an intense laser pulse was envisioned as a method to heat,

vaporize and electronically excite the atomic species present in liquid metal. The

plasma so created would then decay by the emission of photons of frequency

characteristic of the atomic species. By measuring the intensity of particular

8
frequencies in the emitted spectrum, a measurement of the concentration of each

element could be obtained.

In 1983, a cooperative research program began at the DOE's Los Alamos

Laboratory to explore the use of pulsed lasers for the production of representative

plasmas - the first prerequisite for implementation of the laser approach. This work

showed that when pulsed lasers with moderate power density are used to cause

surface ablation, preferential vaporization of volatile elements occurs. This results

in an overabundance of emission from volatile elements in the emission spectrum

and hence a potentially serious inaccuracy in determined composition.

However, researchers at Lehigh University applied plasma transport theory to model the plasma formation process at alloy surfaces and have found that it is theoretically possible to overcome this sampling problem and attain complete

vaporization (of all atomic species regardless of volatility) by operating in a specific

pulse energy density/pulse duration domain. This has led to the recent

implementation of a new program centered at Lehigh University under the

leadership of Professor Y. Kim. The program began in late 1985 and is aiming to develop a mill-worthy prototype within two years.

The budget for this program exceeds $0.5 million and is supported by a Consortium of AISI member companies (including two from Canada), one

aluminum producer and the Ben Franklin Foundation. There is also close

collaboration with a related program underway in the Center for Metals Production

at the Mellon Institute.

ONLINE INSPECTION FOR SURFACE DEFECTS ON HOT AND COLD STRIP

The objective of this program is the inprocess detection, characterization and

classification of surface defects on hot or cold strip steel so that inprocess control can be used to enhance quality. In order to achieve this objective, the sensor would

have to:

• Have a high speed capability commensurate with strip speeds up to 5,000

feet per minute (posing serious problems for data acquisition and analysis

with present commercially available systems).

• Be capable of defect classification and the separation of significant defects

and spurious indications (e.g., rolling solution droplet).

• Have a high reliability in the harsh rolling-mill environment.

It was envisioned that a successfully developed sensor would be used in hot

strip mills, pickle lines, tandem mills, temper mills, and online inspection

operations. Its potential benefits would include:

9
• Early identification of problems in processing which produce defects,

facilitating their prevention at source.

• Elimination of costly reprocessing of defective material.

• Reduction of manpower presently required for visual inspection.

• Elimination of production delays caused by present slow manual

inspection methods and the need for reinspection.

• Improvement of quality of shipped material.

• Provision of extensive documented information in digital form

confirming an overall high level of consistent quality together with the

exact location of isolated defects.

These anticipated benefits were estimated in 1982 to be potentially capable of

annual savings of $50 million in the domestic steel industry. The improvement in

quality and consistency could also be very important factors in the international

marketplace. Furthermore, the sensor was recognized to have potential for similar

inspection and control needs in the nonferrous metals industry.

Technical approaches proposed to the task force unit included the use of eddy

current techniques, ultrasonics and optical scattering. The unit determined that the

optical scattering approach was the most promising for surface inspection. It was

recognized, however, that future subsurface defect inspection needs might

incorporate eddy current and/or ultrasonic technologies.

On July 15, 1984, a contract was signed with the Westinghouse Corporation

for phase I of a three phase sensor development program based on an optical

scattering approach. The initial contract was for a two year study of feasibility and

development of prototype technical specifications. The research budget was entirely

funded by a group of eleven steel and aluminum companies under AISI auspices.

The technical approach being pursued at Westinghouse is a coherent light

scattering approach, shown schematically in Fig. 1. An intense collimated laser

beam is rapidly scanned across the width of the moving metal strip. Detector arrays

are positioned across the strip width at angles predetermined to optimize the defect

scattering (signal) to background scattering (noise) ratio. The detector outputs are digitized and processed, the processed information stored, and the presence of

defects appropriately indicated.

Technical issues that are being addressed include a variety of strategies

designed to maximize the signal-to-noise ratio (contrast) of important defects,

cataloging the characteristic features (signatures) of both defects and spurious

indications for characterization purposes, developing very high speed data

10
acquisition systems, designing and evaluating data reduction schemes and

developing software so that true defects may be discriminated from spurious events

and the defects characterized. The ultimate objective is a full mapping and

characterization of all surface defects in excess of 0.8 mm.

An extensive contrast ratio data base has been compiled for numerous defects

on a wide product range. Techniques being developed to address the defect

recognition problem can be classified as one of three types:

• Implementation of measurement techniques that are minimally sensitive

to false defects.

• Use of additional synchronized measurement systems that locate only the

false defects, for example, drops of rolling solution.

• Development of advanced pattern recognition software that recognizes the

characteristic signatures of defects and classifies them.

Even though the first two approaches are likely to filter out many of the false

defects, some are likely to remain, and advanced software will be necessary to recognize the signatures of these false defects and to classify the true defects.

Progress to date has been sufficiently encouraging that a test facility is being

prepared on a strip mill at the Weirton Steel Corporation Works in Weirton, West

Virginia. This facility will be used to gather information, gain experience of mill

operating conditions, and to perform tests on a range of product containing carefully

documented defects. For this latter purpose, an extensive inventory of

representative defects has been obtained from participating steel and aluminum

companies. This study is presently anticipated to lead to the construction and

installation of a full prototype system for mill evaluation within two years.

RAPID MEASUREMENT OF TEMPERATURE DISTRIBUTION WITHIN A

SOLIDIFIED OR SOLIDIFYING BODY OF HOT STEEL

In order to more effectively control continuous casting and various other

thermomechanical processing treatments (for example, liquid center rolling), it is important to have accurate knowledge of both surface and internal temperature.

The objective of this program is the development of a sensor for the rapid direct

measurement of temperature distributions in both solid and solidifying metal

bodies.

Presently, temperature is determined on the basis of surface temperatures

measured with radiation pyrometers, internal temperatures selectively measured

with thermocouples, or temperatures inferred from process parameters (for

example, rolling pressure)- Computer-based thermal models are also increasingly

being used to predict temperature distributions but with only varying degrees of

11
success. Each has associated limitations and difficulties and alone are unable to satisfy present needs.

An ideal temperature sensor would measure temperature to within

± 10°C with a spatial resolution of < 5 mm over a temperature range of 500° - 1350°C

without need to physically contact the body. It should be capable of giving a reading

within a few seconds and be unaffected by the presence of combustion gas products

or nearby heat sources.

There are numerous potential applications for a temperature sensor during

steel processing. They include:

• Control of cooling during continuous casting.

• Reheat furnace control.

• Hot connection rolling of continuously cast slabs.

« Batch annealing furnace control.

During continuous casting, for example, it would be desirable to mount

sensors just after the strand exits the mold, at several locations in the cooling region,

and before the stand is cut into lengths. This would enable solid shell thickness to be continuously monitored, and allow calculation of the heat transferred through

the mold and during various stages of spray cooling. In turn, this would enable

faster casting speeds, and thus productivity improvements estimated to be ~5%.

Calculations have shown that full implementation of such a sensor in 1982

would have resulted in $275 million savings in energy and productivity

improvements annually for the domestic steel industry.

The task force unit evaluated a variety of candidate measurement

methodologies including ultrasonics, eddy currents, neutron thermalization and

radiography. The ultrasonic approach was deemed in 1983 to have the greatest

likelihood of success although the eddy current approach also had much to

commend it. Discussions were held both with the National Bureau of Standards

(NBS) and the Battelle Pacific Northwest Laboratory (PNL) that ultimately led to complimentary research programs at both laboratories. The program at PNL was co- funded by AISI and the U.S. Department of Energy. The program at NBS was

arranged somewhat differently. NBS assigned its staff to the area while task force

funds were used by the AISI to sponsor the reassignment of steel industry

researchers to NBS to participate in the research program. The programs at both

NBS and PNL also interacted with vendors of high temperature electromagnetic

acoustic transducers (EMAT’s).

12
The ultrasonic approach is based upon the strong dependence of ultrasonic

velocity with temperature (Fig. 2). Each steel has a slightly different dependence, but

for most, the velocity of longitudinal waves generally decreases with temperature at a rate of between 0.5 and 1.0 ms-1 °C-1

. Thus, if the internal velocity distribution

within a hot steel body can be measured, then the temperature distribution can be

inferred using calibration data such as that shown in Fig. 2. The problem of

developing a temperature distribution sensor resolves itself into the separate

problems of reconstructing an ultrasonic velocity profile (or distribution) within a body, obtaining reference data relating ultrasonic velocity to temperature and the

development of ultrasonic instrumentation suitable for reliable high temperature

work in a mill environment.

The latter aspect to the program is presently being addressed by the use of

(remote) high intensity pulsed laser generation of ultrasound coupled with

noncontact EMAT detection. This avoids the need to physically contact the steel.

Such an approach has been demonstrated to be feasible in the laboratory on cooling

steel bodies at temperatures up to ~750°C (Ref. 3). This upper temperature

limitation results from the high temperature degradation of EMAT performance, a limitation that is being addressed in present work at PNL.

The reconstruction of a velocity distribution is being tackled by two

approaches. One is an extension of time-of-flight (TOF) tomography, while the

second is a novel approach that we refer to as dimensional resonance imaging. In

the tomographic approach (Fig. 3), the time-of-flight of ultrasonic pulses along ray

paths of known lengths are used to determine the coefficients of a series

representation of the temperature field. To date, profiles have been determined for

stainless blocks of both cylindrical and square cross section at temperatures up to 750°C (Fig. 4).

A potential concern with the TOF approach is high frequency ultrasonic

scattering by large columnar grains in solidifying bodies and dissipation of high

frequency signal components by temperature dependent internal friction

mechanisms in solidified steel at high temperature. In anticipation of this, a new

low frequency resonance method has also been developed and successfully used to reconstruct temperature profiles in model laboratory experiments. This approach is based upon the shifts in dimensional resonance frequencies from integer multiple

values of the fundamental when a distribution of modulus /density exists within.

An example reconstruction is shown in Fig. 5, where a 1.6°C temperature change

was resolved in a 36" long aluminum bar. Because the frequencies of interest are very low (in the kHz range), grain scattering and internal friction losses are minimized, and sensitive temperature measurements are attainable.

The present objectives are to develop a complete prototype sensor

specification during 1986 and to construct and evaluate the sensor at selective mill

sites during 1987.

13
AUTOMATIC DETECTION OF PIPE AND GROSS POROSITY IN HOT STEEL

BILLETS, BLOOMS OR SLABS

The objective of this sensor development program, more than any other, is changing with time due to a fundamental shift in solidification practice within the

steel industry. At the outset of the program, the majority of steel was solidified by

ingot casting methods and thus pipe and gross porosity were important quantities to locate for control of ingot cropping. With the move to continuous casting, interest

is shifting to the development of a sensor to locate small scale porosity and detect

segregates, defects that are the cause of poor quality during continuous casting. This

has resulted in a continuous need for spatial resolution improvement in the sensor.

The initial aim of the program was development of a sensor to detect primary

pipe and gross porosity in hot, primary-rolled materials in order to provide

sufficient information to determine an optimum cropping point. The sensor was

expected to:

• Locate the edge of pipe to within 1/10" slabs up to 12 inches thick.

• Detect and locate pores greater than 1/3" in diameter.

• Perform a two dimension scan within 30 s or less.

• Be capable of operation at steel surface temperatures of 830°-1100°F.

The availability of such a sensor would benefit productivity and yield

through the avoidance of discarding sound steel. In 1982, yield improvements of as

little as 0.5% were considered capable of generating a $50 million benefit to the

domestic industry annually. Other benefits would accrue from an improved quality

through avoidance of pipe and porosity in finished material. This would help

counter foreign competition, lead to greater customer satisfaction, and result in

significant energy savings by eliminating the need to cool to ambient temperature to perform internal (ultrasonic) inspections.

Past research had shown ultrasonic scattering to be capable of detecting defects

of the size of interest when the steel was at ambient temperature (where it is simple

to get good ultrasonic coupling). On hot products, however, it is necessary to

thermally protect the sensitive piezoelectric materials used to generate and detect

ultrasound. This has been achieved in the past through the use of cooled metal bars

(buffer rods). One (cooled) end is attached to a piezoelectric transducer, and the

other pressed against the body to be inspected (Fig. 6). While the approach was

sufficiently sensitive, very large pressures were found necessary in order to couple

the ultrasound into the steel body (Ref. 4). This was found to cause the system to be

mechanically unreliable and very slow. The adoption of noncontact ultrasonic

approaches similar to those proposed for a temperature sensor therefore appears to

14
be a promising route to a solution, though one not expected to be realizable in the

short term, given the still developing state of noncontact ultrasound technology.

The pressing need for a short term, at least partial solution, to the problem

resulted in the evolution of a dual-thrust collaborative program involving the

AISI, Magnasonics*, Magnaflux*, NBS and more recently the Argonne National

Laboratory (ANL). The short term (interim) solution was to be addressed by

evaluation of a rolling wheel contact method proposed by the Magnaflux

Corporation. The approach featured a rolling wheel (in place of a buffer rod) that

promised to overcome the scan speed limitation. The longer term solution was

perceived to be the development of high temperature EMAT's and (possibly) pulsed

laser excitation of ultrasound. From a technical viewpoint, the early

implementation of a contact method was also perceived to be a useful means for

gathering important information about the ultrasonic scattering properties of

realistic defects and the nature of wave propagation in very hot (as-cast

microstructure) steels. Thus information would be important in guiding the longer

term noncontact approach.

Research to date has shown the contact wheel approach to be capable of the

required resolution and speed at ambient temperature where vacuum grease

couplants can be used to eliminate the need for high pressure. However, at higher

temperatures, these particular couplants do not work, and high pressure and

alternative couplants based upon various salts, glasses and fluxes have been

investigated in research at NBS and Magnaflux. One couplant in particular, borax,

has been found by NBS to give excellent ultrasonic coupling, even at 1100-1200°C . For example, in Fig. 7, we show the signal from a pitch-catch ultrasonic

configuration, (shown in Fig. 6) in one case at ambient temperature using a vacuum

grease couplant, and in the other at 800°C using borax. Given that ultrasonic

propagation is usually worse in steel at high temperature, the better signal-to-noise

ratio of the defects signifies a very high coupling efficiency, sufficient to eliminate

all need for additional pressure coupling.

Prior to installing a prototype rolling wheel system in a mill, full scale trials are being conducted at the ANL where heating/handling facilities exist for large

steel bodies. Encouraging progress has been made to date, and plans are progressing

for mill evaluation of a prototype within one year.

Encouraging progress has also been made with the noncontact approach.

Electromagnetic acoustic transducers have been designed for both ultrasound

generation and detection with optimum sensitivity for the pipe/ porosity detection

at high temperatures. They have been evaluated at temperature on ferritic steels

both above and below the Curie temperature. As noncontact generators of

ultrasound, EMAT's to date are not as efficient as would be desired, but

improvements are being pursued. As receivers of laser excited ultrasonic pulses,

considerable early success has been achieved, encouraging us to believe that a noncontact solution may be viable within a few years. Work is also continuing to

15
optimize the instrumentation, improve its survivability at high temperature,

develop signal processing software and to improve spatial resolution for the

detection of small defects.

THE FUTURE

The collaborative research programs initiated by the AISI are each now

showing promise of at least partially solving the most immediate sensor needs of

the steel industry. As encouraging as this research progress is, the ultimate test of its usefulness will be the industry implementation of these process control tools. This

may be as difficult to achieve as the research has been to perform because it raises

several, as yet unresolved, issues. For instance who will build and install these

sophisticated sensors in steel mills? Surely this is not a task suitable for the federal

laboratories involved in much of the research nor the steel companies themselves,

who no longer retain the necessary in-house expertise. The most promising

approach at this time appears to be the assignment of the technology to vendor

companies under some form of license arrangement that protects the vendors

subsequent investment. The emergence of an economically healthy, state-of-the-art sensor vendor industry therefore seems to be an important factor in determining

the future availability of process control sensors for steel. The industry must

therefore now look to ways to promote this important development.

The collaborative nature of the research programs described above arose because of a shared need for sensors by the entire domestic steel industry, combined

with their high development cost and risk (prohibitively high for a single

company). These factors are seen increasingly in other steel research areas, for

example, in the development of new processing technologies such as thin strip

casting. Here also strong collaborative programs are now emerging with the intent

of advancing the competitivity of the domestic industry in the international

marketplace. It seems that collaborative efforts are on the increase, and those for sensors can form a model for this approach to meeting the research needs of an

industry. The careful ways in which research needs were identified, ongoing

research monitored, and results disseminated provide good guidelines for all collaborative programs. The reassignment of steel industry staff to the laboratories

pursuing research has been a particularly effective two-way technology transfer

mechanism.

At the beginning of the programs, sensors, were viewed as inprocess

inspection tools or as useful adjuncts to existing process control strategies.

However, recently, the very exciting opportunity of a new process control strategy,

the so-called Intelligent Processing of Materials, is now beginning to emerge handin-hand with the emergence of these sensors (Ref. 5). Intelligent process control

involves three elements:

16
• Predictive process models.

• Sensors.

• Control systems featuring artificial intelligence /expert systems.

For example, sophisticated models for alloy solidification have been

developed in recent years and are presently used to define operational conditions

during continuous casting. These models need, as inputs, information about key

process and microstructure variables (temperature gradient, solidification interface

velocity, interface morphology, heat transfer rates, etc.). Presently estimated values

are used and this introduces substantial uncertainties. The development of sensors

able to measure these quantities inprocess greatly expands the usefulness of the

process models, for now they can make detailed accurate predictions of the current

state of solidification. Adjusting control parameters (such as heat extraction rate,

mold lubricity, casting rate, etc.) then provides a means for continuously

maintaining optimal solidification conditions consistent with productivity

requirements. However, with numerous (nonlinear) interacting variables to be

controlled in rapid time frames, automated control schemes are required. It is the

recent development of artificial intelligence/expert systems that offers such great

promise for such control schemes.

Intelligent processing of materials is consistent with the implementation of

full factory automation systems and with them, the opportunity to attain substantial

productivity and quality improvements. The development of intelligent processing

strategies may also expose opportunities to implement commercially entirely new

methods of processing -- methods hitherto considered too unstable for commercial

utilization -- a development that would truly advance the competiveness of

domestic materials producing industries.

ACKNOWLEDGEMENTS

In preparing this report, I have drawn heavily upon unpublished

research prospectuses developed by AISI Task Force Units. I am also grateful

to the AISI for permission for this. I am grateful to Drs. Y. Kim, T. Porter,

F. Achey, D. Rogers, J. Cook, W. Wilson, B. Sylvester, K. Kappmeyer, R.D.

Jeffress and W. E. Dennis who have provided most helpful discussions on

numerous aspects of the work reported.

Disclaimer: Individual companies are mentioned in this paper for the purpose of

accurately representing the research groups involved in collaborative programs.

This should not be construed as an NBS endorsement of their products.

17
REFERENCES

1. Steel Industry Research Priorities for Process Control and Sensor Development.

1981, AISI Internal Report.

2. R, Mehrabian, R.L. Whiteley, E.C. van Reuth and H.N.G. Wadley, 1982, Process

Control Sensors for the Steel Industry: Workshop Report, National Bureau of

Standards Internal Report 82-2618.

3. H.N.G. Wadley, S.J. Norton, F. Mauer and B. Droney, Ultrasonic Measurement of

Internal Temperature Distribution, Proc. Roy. Soc. (London), 1986, In Press.

4. B.E. Droney and T. J. Pfeiffer, Ultrasonic Inspection of Hot Steel Blooms to detect

Internal Pipe, Mat. Eval. June 1980, p. 31.

5. R. Mehrabian and H.N.G. Wadley, Needs for Process Control in Advanced

Processing of Materials, Journal of Metals, Feb. 1985, p. 51.

18
Fig. 1 A photodiode array is used to detect the coherent light scattered by defects

on moving metal strip.

5800

TEMPERATURE, *C

Fig. 2 The dependence of ultrasonic velocity upon temperature for AISI 304

stainless steel. The insert shows the experimental configuration used

to measure ultrasonic time-of-flight between points A and B.

19
Fig. 3 An intense laser pulse is used to induce a strong ultrasonic pulse

whose time-of-flight over various ray paths is measured using an

EMAT and associated high speed digital instrumentation.

NORMALIZED DISTANCE

Fig. 4. A reconstructed temperature profile through the thickness of a six inch

thick 304 stainless steel block (curve). Also shown are simultaneously

recorded thermocouple measured temperatures for comparison.

20
Measurement of Internal Temperature

Distribution

Reconstruction From Five Dimensional

Resonance

Ambient + 1.6°C

<9

b.3«

<sQ.E

«

Ambient

Aluminum

I

--

kZ& Furnace Zone

E^i

18 36

Distance, Z/inches

Fig. 5. A reconstructued temperature profile using the dimensional

resonance technique.

PROBE ASSEMBLY

Fig. 6. Cooled steel bars can be used to thermally isolate ultrasonic

transducers.

21
COLD

TEST90

TO

NORM.

ROUGH.

SIDE

FLAW SIGNALS

COLD STEEL HOT STEEL

OE rECT

l

I

1 1

0.00 60.00 120.00 HICROSECONOS

Fig. 7. Ultrasonic waveforms showing detection of defect echos at ambient

and 800°C.
RECENT PROGRESS IN THE EVALUATION OF A PIPE/POROSITY SENSOR

DAVID KUPPERMAN

Argonne National Laboratory

Argonne, IL 60439

INTRODUCTION

In primary rolled ingot steel, a variety of both large and small

imperfections occur at the end of the rolled slab, and must be sheared off prior

to conversion into a final product. Presently, visual observations on steel at 2000°F are used to estimate the optimum amount of material removed by

cropping. These observations are insufficient to determine the point within

the steel where flaws end and "sound" steel begins. Thus, overconservative

cropping is often practiced that leads to $50 million losses in the domestic

industry annually.

To help alleviate this problem, a new device, that can detect

imperfections as small as 0.25" within hot steel slabs, has been developed

during a two-year collaborative effort between AISI member companies, NBS,

and Magnaflux Corporation. The new development, a "high-temperature

imperfection detector" uses ultrasonic pulses to detect internal discontinuities

in hot steel. In operation, a transducer capable of generating ultrasonic

pulses, is in contact with the inside rim of a three-foot diameter stainless steel

wheel (Fig. 1). This wheel rotates as steel slabs at 2000°F pass underneath.

The transmitted ultrasonic pulses transmitted into the steel through the rim

are reflected by internal discontinuities. These reflected signals pass back into

the wheel and are detected by a receiver transducer. To evaluate the device, a hot testing facility has been assembled at the Argonne National Laboratory

(ANL) and tests conducted upon numerous samples containing naturally

occurring defects representative of those encountered in a primary rolling

mill.

PROTOTYPE EVALUATION

The prototype device described above is augmented by a computer

based analysis system that records the amplitude and time-of-flight of

reflected pulses. The system then produces a color map of the internal defect

structure. Each imperfection is outlined according to size and location and is color coded based upon the amplitude of the detected ultrasonic reflection.

The ultrasonic system has been designed to scan representative 8" x 18" x 20" slab sections that are near the cropping temperature of 2000°F. The

configuration and slower speed of this prototype depart somewhat from that

23
<

Fig. 1 A photograph of the contact ultrasonic sensor undergoing evaluation

upon a hot steel billet at the Argonne National Laboratory.

24
which eventually will be required for production installations; however, the use of multiple transducers and a faster computer will overcome these

problems.

For the thick steel slabs evaluated, a one-sided "pitch-catch" split wheel

was selected. However, the system configuration, size, and approach could be

conveniently changed for billet, bar, tube, and strip applications (Fig. 2).

Fig. 2 The split wheel ultrasonic sensor developed for detection of gross

porosity and pipe.

Ultrasonic coupling is a major problem. Work carried out at Argonne

led to the following conclusions; either of the following three methods could

be applied for coupling of the ultrasonic waves to the hot steel: (a) high

pressure direct coupling; (b) magnesium oxide powder with moderate

pressure; or (c) melted borax with very minimal pressure. To ensure longterm system mechanically reliability, the melted borax approach may be

preferred. In the event the borax is applied, a zirconium or other protective

coating would be applied to sections or the stainless steel wheel to avoid the

corrosive effects of borax.

m

25
Several comparison between cold ultrasonic immersion tank images

and the color images from hot testing provided progressively improved

correlation as the testing's program was carried out Eight-inch crop ends,

containing primary pipe, fishtail, and secondary pipe, were scanned with the

wheel in both cold and hot states. In addition, system repeatability was

checked and correlation was excellent. The adequacy of the thermal analysis

was also demonstrated by successfully predicting resulting temperature

changes in the wheel and ultrasonic probe.

The most recent tests of the sensor system at Argonne (including a successful demonstration to representatives of 24 steel companies) have

confirmed that such a device could be used inline within a steel mill

environment to clearly define the imperfections in hot steel. A defect as small as 1/4" in diameter can be detected at 2000°F (Ref. 2). The widespread

use of this concept in steel rolling and continuous castings could result in

considerable monetary savings that would improve the American steel

industry competitiveness throughout the world.

REFERENCES

1. CD. Rogers, B.E. Droney, J.M. Toth and H.N.G. Wadley "Detection of

Pipe and Porosity in Hot Steel Slabs, Blooms and Billets", Proc. 26th

Mechanical Working and Steel Processing Conference, October 1984.

2. D. Kupperman, "Design and Testing of a Prototype Contact Ultrasonic

Sensor System for Detecting Gross Porosity within Thick Hot Steel

Slabs, Blooms and Billets", In preparation.

26
A PULSED LASER/ELECTROMAGNETIC ACOUSTIC TRANSDUCER

APPROACH TO ULTRASONIC SENSOR NEEDS FOR STEEL PROCESSING

G.A. ALERS

Magnasonics, Inc. 215 Sierra Drive, SE

Albuquerque, NM 87108

INTRODUCTION

Many of the traditional NDE techniques of the past are today being

investigated for their potential role as process control sensors for materials

processing (Ref. 1). Ultrasonic inspection appears one of the most promising

because of its ability to penetrate opaque bodies and to allow determination of

microstructure variables (such as grain size), process variables (such as internal temperature distribution) and detect internal discontinuities (cracks,

pores and inclusions). A key problem with the traditional approaches to

ultrasonic measurements is the need to contact the body being probed with

piezoelectric transducers. These transducers are fragile, require couplants,

and fail when exposed to temperatures of more than a few hundred °C. Their

use during processing may thus require practices that unacceptably interfere

with the process.

Pulsed lasers for remote ultrasonic generation have been available for

several years. Their use would overcome all the difficulties above if a similar

remote ultrasonic receiver were available. The obvious candidate for this

would be a laser interferometer, but it is of insufficient sensitivity for

applications which involve unpolished, poorly reflecting surfaces. An

alternative noncontact (but no longer remote) receiver could be an

electromagnetic acoustic transducer (EMAT). This has the potential to be

engineered to survive the process environment, and if carefully designed, to

match the characteristics of a laser ultrasonic source.

Electromagnetic Acoustic Transducers (EMATs) are noncontact

ultrasonic transducers that can operate on metals across a small air gap. Thus,

they should be applicable to the problem of defect detection in steel at very

high temperatures. During the past few years, the AISI and NBS have

sponsored an advanced development program aimed at demonstrating the

proof-of-principle for combining an EMAT ultrasonic receiver with a pulsed

laser generator of ultrasonic pulses. This report summarizes the results

obtained to date and includes an addendum describing very recent

developments sponsored by the DOE and Battelle Northwest Laboratories.

27
LASER GENERATED ULTRASOUND

The generation and subsequent propagation of ultrasound by the

absorption of an intense laser pulse is now a relatively well understood

process (Ref. 2). Research at Hull University and Harwell in England during

the early 1980's indicated that two modes of laser generated ultrasound may

occur. For low absorbed fluxes, the surface where absorption occurs never

exceeds its melting temperature and the source of ultrasound is then a transient dilatation. The stresses associated with this dilatation are for the

most part below the elastic limit, and this mode of generation is therefore

referred to as thermoelastic. At higher fluxes, the surface temperature rise is capable of exceeding the vaporization temperature. Atoms leave the surface

at high velocity imparting a momentum to the substrate that is the source of

ultrasound. This mode of generation is referred to as ablation.

It is possible to model the thermoeleastic source using heat flow theory.

For simple bodies, such as isotropic elastic half-spaces and infinite plates, this

theoretical description of the source can be coupled with the bodies dynamic

elastic Green’s tensor of the bodies to quantitatively predict the temporal

waveform of the ultrasonic signal (Ref. 2-4). The displacement signals at epicenter and various locations on the plate surface are shown in Ref. 3. The

response of an EMAT is proportional to surface velocity and so the velocity at

the longitudinal arrival at epicenter is a positive pulse whose half width is approximately that of the source duration (in the absence of grain /scattering

absorption) while at the shear arrival it is a negative pulse.

The ablative source is not so amenable to quantitative modelling

because we do not know the mass or velocity of material ablated, or how this

varies with time during optical absorption. However, if it is assumed that the source is mechanically equivalent to a momentum of arbitrary strength,

temporal waveforms of arbitrary amplitude may then readily be evaluated

(Ref. 2). As seen by an EMAT, the principle difference from a thermoelastic source is that the epicenter longitudinal arrival is now a bipolar pulse with an

initially negative direction. For this source, the epicenter shear arrival

remains a unipolar pulse but of opposite sign to that of the thermoelastic

source.

In practice, it is not possible to produce pure ablation without

accompanying thermoelastic contributions to the source. The relative

contributions of each to the emitted ultrasound are yet to be fully resolved,

and will vary with optical flux. However, the interesting possibility exists of

identifying a range where only longitudinal waves propagate because of

cancellation of the shear components due to thermoelastic and ablative sources.

28
The displacement amplitude of a laser generated ultrasonic signal is proportional to the absorbed optical energy in the thermoelastic region. Thus,

the velocity amplitude will be proportional to the absorbed optical power.

This should be maximized if we wish to maximize the EMAT signahnoise

ratio. At NBS, a Q-switched Nd:YAG laser is available with a 25 ns pulse

duration and variable energy per pulse up to 800 mj (average power of 32

MW). 175 mj, 1 mm diameter pulses were used for the results reported here.

EMAT ULTRASONIC RECEIVERS

Although the EMAT is technically a noncontact ultrasonic transducer,

it must be held close to the surface before the electromagnetic induction

process can operate with a useful efficiency. For application to inspection of

hot steel, any small air gap at the surface of the metal acts as a good thermal

insulator and greatly reduces the heat transfer into the transducer structure.

In addition to being physically capable of operating in a high temperature

environment, the EMAT has several other features that can be exploited to advantage. By simply changing the direction of the magnetic field used, the

transducer can be made to detect either longitudinal or shear waves. Also, by

choosing the shape of the coil of wire inside the EMAT, the angle of incidence

of the waves to which it is sensitive can be controlled. Since the electromagnetic induction process can be made to operate over a wide range of

frequencies, EMATs are fundamentally broadband devices. Their frequency

of operation is usually set by the electronic receiver circuits used to amplify

the electrical signals. For hot steel, the best frequency of operation is determined primarily by the laser source spectrum, the ultrasonic attenuation

properties of the steel and only secondarily by the structure of the EMAT and

its associated electronic circuits.

An important problem to be overcome when designing an EMAT for

operation on hot steel is to be sure that the electrical resistivity of the hot steel

is not so high that the electromagnetic skin depth becomes comparable with

the wavelength of the ultrasonic waves involved. For ordinary metals with

resistivities near 10 (iohm-cm, this criterion is easily satisfied for frequencies

less than 10 MHz. However, hot steel has a resistivity between 100 and 200

jiohm-cm so the frequency response may be attenuated above 1 MHz for

longitudinal waves and 0.3 MHz for shear waves.

EMAT DESIGN

Since an EMAT consists of a coil of wire held close to the surface of the

part being inspected plus a large magnetic field to flood the region around the

coil, two problems must be addressed to enable hot steel bodies to be interrogated. First, the coil must be constructed with high melting point wires and

a layer of thermal insulation between the coil and the steel. Second, the

29
magnet must be designed to supply a large magnetic field to the total area of

the EMAT coil and must be protected from the heat generated by the hot steel.

The most direct approach to supplying a magnetic field to the EMAT

coil is to use a large electromagnet with pole pieces designed to concentrate

the magnetic flux at the EMAT coils, as shown in Fig. 1(a). Here, the EMAT

coils can be located in either a normal or tangential field direction and hence

to receive either longitudinal or shear waves. (Position A in a tangential field

is suitable for longitudinal waves while position B, in a normal field, is suitable for shear waves.) Such an electromagnet was constructed but proved

to be quite massive and difficult to move in a scanning apparatus even

though it generated magnetic fields in the 4 to 5 koe range at an EMAT coil

with a 1/2" diameter. Figures 1(b) and (c) show magnet configurations based

on samarium cobalt permanent magnets located very close to the EMAT coil

itself in order to maximize the fields at the coil wires. Both of these structures

were assembled out of 1/4" thick by 1/2" square slabs of SmCo and fields in

the range of 3 to 4 koe were measured at the EMAT coil positions. These

permanent magnet structures were quite light and maneuverable and thus

could be easily scanned over large areas. Their reduction in efficiency caused

by slightly smaller magnetic fields could be tolerated in exchange for this

simplification in mechanical structure.

The EMAT coils were made thermally stable by housing their copper

wires in small (1/16" diameter), four-hole ceramic thermocouple tubes. Since

the duration of the exposure to high temperature was intended to be short in

order to protect the permanent magnets, it was deemed unnecessary to use more expensive, high resistance wires such as platinum for the EMAT coils.

By using thermocouples tubes with 1/16" diameters, the minimum liftoff or separation between the EMAT coil and the sample was 1/32" (0.8 mm). In the

regions at the ends of the thermocouple tubes where the copper wires were

unprotected, the wires were coated with a high temperature ceramic cement

that served to thermally and electrically insulate each wire from its surroundings.

Since it was of interest to make time-of-flight measurements and since

the attenuation of high frequency shear waves in hot steel is higher than

longitudinal waves (Ref. 5), the high temperature EMAT designed to detect

longitudinal waves propagating directly through the thickness dimension of

the steel samples received the most attention. Thus, the comb magnet

construction technique shown in Fig. 1(c) was used in all the experiments

reported here. All the wires in each section of the EMAT coil between the

pole pieces conducted current in the same direction but this direction

reversed in the adjacent pole piece gap where the magnetic field was also in

the opposite direction. In this way, the entire face of the EMAT would

respond in phase to a plane longitudinal wave striking the face along its

30
(a)

Fig. 1 Magnet configuration needed to operate EMATs. (a) Electromagnet for

exciting or detecting either longitudinal or shear waves, (b) Permanent

magnet arrangement for exciting or detecting shear waves, (c) Permanent magnet design for excitation or detection of longitudinal waves.

normal direction. This method of assembling the permanent magnet

structure allows the pole pieces to conduct heat directly to the samarium

cobalt. Thus, the ability of the probe to operate on hot objects depended

critically on the length of time it took to heat the permanent magnets to a temperature at which their performance was jeopardized. The manufacturer

of the magnets used in these EMATs set the upper temperature limit for his

product at 330°C for continuous operation. During the tests on hot steel

blocks, a thermocouple on the samarium cobalt showed that it required over

1-1/2 minutes for the temperature of the magnets to exceed 250°C. This 90 s interval was ample time for collecting all the necessary ultrasonic

information. Therefore, it was not necessary to add cumbersome water

cooling apparatus at this stage.

31
RESULTS

In accord with the ablative mechanism of generating ultrasonic waves,

very large ultrasonic signals were observed with the longitudinal wave

EMAT shown in Fig. 1(c) when it was positioned on the face of a sample

directly opposite to the point at which the laser beam struck the sample.

When a shear wave sensitive EMAT was used (Fig. 1(b)), shear waves were

detected only when the EMAT was displaced from epicenter along the back

side of the sample. As expected from EMAT theory (Ref. 6), shear waves at an

angle of 30° relative to the surface normal were easily detected. One

surprising result was the observation of a second large shear wave signal

leaving the surface of optical impact at an angle of 60° relative to the surface

normal. This may be a manifestation of the thermoelastic contribution to the source.A very important performance characteristic of the longitudinal wave

EMAT used here was the amount of air gap or liftoff that could be tolerated

between the sample and the front face of the EMAT. Figure 2 shows how the

output signal from the EMAT decreased as a function of this separation

distance. This graph shows that the addition of up to 0.02" of thermally

insulating material between the pole pieces and the hot object would not cause a significant reduction in sensitivity but would probably make a dramatic improvement in the ability of the structure to withstand exposure to high temperatures. The fact that the drop in signal strength is described by an

exponential function is consistent with the periodic structure of this form of

magnet. The rate of drop could be decreased (to achieve higher sensitivity at

large air gaps) by using thicker permanent magnets to increase the separation

distance between pole pieces.

Fig. 2 Dependence of the EMAT receiver probe sensitivity on the thickness of

the air gap between the probe and the surface of the sample.

32
A second important performance characteristic of the longitudinal

wave EMAT used here was the frequency bandwidth that could be achieved

with that specific EMAT design and amplifier circuit. Figure 3 shows the

waveforms observed when the longitudinal wave EMAT shown in Fig. 1(c)

was used to receive directly transmitted longitudinal waves generated by the

pulsed laser. When narrow band filters were used in the amplifier stages, the

waveform shown in Fig. 3(a) was observed. Note that the noise received

prior to the ultrasonic wave arrival is very small and the electrical signal is a tone burst containing many cycles of oscillation. Figure 3(b) shows the effect

of removing all the filters and tuning capacitors from across the EMAT coils

to achieve a large bandwidth. The noise level prior to the arrival of the

acoustic wave is much higher than in the previous cases. However, it appears that the arrival time could be measured to an accuracy greater than

0.1 (is (or 0.6%).

0.50

0.30

0.10

- 0.10

- 0.30

- 0.50

0.50

0.30

0.10

- 0.10

- 0.30

- 0.50

0.00 20.00 40.00 60.00 80.00 100.00 120.00 140.00 160.00 TIME/hs

Fig. 3 Ultrasonic longitudinal waveforms observed with different electronic

tuning circuits connected to the EMAT coil, (a) Band pass filter

combined with a tuning capacitor across the EMAT coil, (b) No

tuning.

33
In order to demonstrate laser excitation and EMAT reception of

ultrasonic waves in hot steel, two large steel block samples were heated in a large furnace at 980°C (1800°F). One sample was made of 304 stainless steel

and hence underwent no ferromagnetic transition as it cooled off. This

sample was a 4" diameter cylinder in which the sound wave propagated along

the 4" long axis of the cylinder. The second sample was a 6" by 6" square slab

of 1018 steel arranged so that the sound wave propagated parallel to the 4"

thickness dimension. For this sample, a transition from the nonmagnetic to magnetic state would be expected when the temperature passed through the

Curie temperature of 770°C (1418°F). These hot blocks were positioned such

that the focal point of the laser beam was at the center of the front surface of

the sample. The EMAT was then rested lightly against the back surface

directly opposite the laser impact point. Thermocouples were held against

the sample surface and inserted into the EMAT structure at the location of the

samarium cobalt permanent magnets in order to monitor the local

temperatures while the steel cooled and the EMAT heated up.

Figure 4 shows examples of the waveform observed on the 4” long,

stainless steel cylinder as its surface temperature fell from 752°C to 322°C. The

directly transmitted longitudinal wave signal and three reverberations can be

easily distinguished from the background noise. Most of this noise is probably from ultrasonic signals reflected by the side walls because the noise is at the EMAT frequency and increases with time after the laser impact. Note

that the noise on the time interval immediately following the laser pulse and

prior to the arrival of the first, longitudinal wave, signal is very low at all temperatures. As the sample became more cool, the acoustic noise between

the reverberation signals increased. This may be explained by mode

conversion at the side walls plus a dramatic lowering of the attenuation for

shear waves formed by mode conversion as the temperature of the steel

became lower.

Figure 5 shows the waveforms observed on the 4" thick slab of 1018

steel as the surface temperature dropped from 810°C (1490°F) through the

Curie temperature to 618°C (1144°F). At the highest temperature studied, the

ultrasonic longitudinal wave signals were very well defined and appear

similar to the signals observed on stainless steel. Below the Curie

temperature, the magnetic field from the pole pieces of the comb-type EMAT

flowed directly into the steel in the immediate vicinity of the pole piece so

that the tangential field at the EMAT coil became greatly reduced. Therefore,

the sensitivity to longitudinal waves fell and the directly transmitted

longitudinal wave which should arrive at about 20 jis was no longer the

dominant signal observed. Instead, a late arriving pair of signals appear,

which are probably shear waves that have reached the EMAT by reflecting

from the side walls of the sample. Since these shear waves have reflected

34
NORMALIZED

EMAT

VOLTAGE

0.50

0.30

0.10

- 0.10

-0.30

-0.50

TIMERS

Fig. 4 Waveforms on a 4" thick stainless steel block.

35
NORMALIZED

EMAT

VOLTAGE

0.50

0.30

0.10

= 0.10

= 0.30

= 0.50

(a)

'

1 4

41

...

i:

jl

f

Average Surface Temperature 810°C

:

i $ f ! I:

1 1 1 •*•»•<»

t •> '.. I

' -

•

1 J-. !- la

: 4' --4 V, . vJ./.4;;44ui4.u4

;k t' . y. ; ; )]| • :i

Fig. 5 Waveforms on a 4" thick 1018 block.

36
from the side of the sample, they must impinge on the EMAT at an angle and

would be detected if a periodic normal magnetic field existed in the EMAT

structure. Such a field is actually present in the immediate vicinity of the

pole pieces of the comb-type EMAT when it is in contact with a sample that is ferromagnetic. An analysis of the geometrical dimensions of the EMAT used

in these experiments indicates that the EMAT would be sensitive to shear

waves that approach the face of the probe at an angle of about 48° relative to

the surface normal. Such an angle is reasonably consistent with the sound

wave paths reflected by the side walls of the sample.

CONCLUSIONS

1. A lightweight, easily scanned EMAT receiver probe has been

constructed out of heat resistant materials and samarium cobalt permanent

magnets to withstand intermittent use at high temperatures. It has been

optimized for use with a pulsed laser ultrasonic source.

2. This probe was able to detect 1 MHz ultrasonic pulses generated by a focused 175 mj laser pulse with a signal-to-noise ratio of 40:1 (32 dB) on a mild

steel block with a surface temperature of 810°C (1490°F). Similar results were

also obtained on a stainless steel sample.

3. No degradation of performance was observed after the probe had

been repeatedly held in contact with hot steel for time intervals that ranged

from 60 to 90 s for each contact even though no active water cooling was

employed. It now appears relatively straightforward to develop EMAT's for

semicontinuous use at temperatures up to 1300°C.

ADDENDUM

Since this AISI/NBS program was completed, Battelle Northwest

Laboratories contracted Magnasonics to assemble an improved EMAT for

measuring the transit time of ultrasonic waves through 9" thick sections of

steel. Here, again, a pulsed laser was used as the ultrasonic transmitter but

the improved EMAT receiver used a pulsed electromagnet instead of

samarium-cobalt permanent magnets and was capable of continuous

operation next to the hot steel because it was housed in a water cooled,

ceramic housing. Preliminary results show that this EMAT put out broad

bandwidth signals with 20 dB signal-to-noise ratios on 9" thick stainless steel

samples with surface temperatures of 2000°F. Signals were still observable

when the EMAT air gap was over 1/4".

ACKNOWLEDGEMENTS

The help and assistance of F. Mauer and J. Martinez (NBS) are

gratefully acknowledged.

37
REFERENCES

1. H. N. G. Wadley, J. of Metals, Oct. 1986.

2. C. B. Scruby, R. J. Dewhurst, D. A. Hutchins, and S. B. Palmer, Research

Techniques in Nondestructive Testing V. ed., R. S. Sharpe, p. 281,

Academic Press, Oxford (1982).

3. H. N. G. Wadley, C. K. Stockton, J. A. Simmons, M. Rosen, S. D. Ridder,

and R. Mehrabian, Review of Progress in Quantitative NDE 1. p. 421,

Plenum Press, New York (1982).

4. H. N. G. Wadley, J. A. Simmons, and C. Turner, Review of Progress in

Quantitative NDE 3B, p. 683, Plenum Press, New York (1984).

5. E, P. Papadakis, et al., I. Acoust. Soc. Am. 52. p. 855 (1972).

6. C. F. Vasile and R. B. Thompson, Ultrasonics Symposium Proceedings.

IEEE Cat. No. 77CH1264-ISU, p. 84 (1977).

38
NONCONTACT LASER ULTRASONIC SYSTEM FOR INSPECTION OF HOT

STEEL BODIES DURING PROCESSING

MOSHE ROSEN AND ROBERT E. GREEN, JR.

Center for Nondestructive Evaluation

and

Materials Science and Engineering Department

The Johns Hopkins University

Baltimore, MD 21218

ABSTRACT

The objective of this paper is to demonstrate the capability of a noncontact laser ultrasonic generation and detection system for detection of

primary pipe and gross porosity in hot steel slabs, blooms, and billets and to

map the location of the defects during processing. Ultrasonic pulses are to be

generated on one surface of a hot steel slab by impingement of a single-pulse

from a high energy laser. The ultrasonic pulses propagating in the slab can be

detected on both the same side as the generator (pulse-echo) and on the

opposite side (through-transmission) by means of a laser interferometer

incorporating a helium-cadmium laser. Assembly of the generation and

detection subsystems on a scanning device permits rapid monitoring of the

desired volume of the hot slab. The approach also affords the possibility of

acquiring information useful for material characterization and processing.

INTRODUCTION

Pipe, or shrinkage cavity is a result of the present practice of the

steelmaking process (Ref. 1). As the molten steel solidifies within an ingot

mold, evolution of gases and thermal contraction during solidification cause

the formation of cavities primarily at the top of the ingot. As the internal

surfaces of the cavity are exposed to the atmosphere, oxide layers are formed.

Upon subsequent deformation of the ingot, e.g., rolling, the oxide coating of

the cavity prevents its welding. These internal discontinuities, of a permanent nature, are highly detrimental to the quality of the rolled

material, and must be cropped off. The cropping operation is presently a manually controlled process. The operator visually inspects the steel slab and

determines the appropriate cropping location based on his acquired

experience. Quite often, the cropping operator shears off several slices before

he is satisfied that all pipe has been removed from the top of the ingot. Also,

in some cases an overly conservative operator shears off excessive good

material in order to assure removal of the internal discontinuities. Such a procedure has proven to be costly and time consuming. It is believed by

technical executives of the steel industry in the U.S. and worldwide that the

cropping operation would be optimized if the operator could receive

39
information from an automatic inspection system capable of detecting and

locating pipe and gross porosity within the slab or bloom. This would render

the cropping process more efficient, reliable and independent of human error

(Ref. 2). An automatic detection and location system would enhance

productivity by diminishing labor and energy intensive operations, and by

improving the quality assurance of the final product.

A substantial investigative effort, worldwide, was devoted to solve this

problem (Ref. 3-11). A wide variety of techniques was examined for

feasibility, e.g., online x-ray radiography, gamma-ray tomography, eddy

currents, electromagnetic ultrasound transduction (EMATS) and contact

ultrasonic methods. In Europe, USSR, Japan and France, and to seme degree

in the USA, the main effort was addressed toward conventional ultrasonic

techniques whereby an array of ultrasonic transducers is subjected to a momentary forced contact with the hot slab. Water spray precedes the

contact, thus affording an acoustic coupling between the transducers and the

steel slab and localized cooling of the slab in order to prevent overheating of

the ultrasonic transducers. In spite of the substantial technological effort, the

conventional bulk ultrasonics technique has met with limited success only

on a laboratory scale. The pressure-coupled ultrasonic transducers have

never been applied for full production. The major steel companies in the

world are actively engaged in searching for a more efficient means to assess the internal soundness of the material (Ref. 2).

The purpose of this paper is to describe a sensing system for online

monitoring and control of the internal soundness of steel slabs and blooms at

elevated temperatures by means of noncontact ultrasonic techniques

employing laser devices for both generation and detection of ultrasonic

waves.

LASER GENERATION OF ULTRASOUND

Application of a transient load on the surface of a metal by means of

rapid energy deposition from a single-pulse Q-switched high energy laser

produces a thermoelastic compressive stress wave that propagates through

the material. Rapid deposition of laser energy into absorbing material

produces a state of unbalanced compressive thermal stress as the heated

materials wants to expand but is restrained initially by its own inertia. The

thermoelastic mechanism for generation of acoustic waves is based on the

solution of the wave equation of forced oscillations of a material under the

effect of a thermoelastic force. The amplitude of the acoustic oscillation is given by the generalized equation of thermoacoustics (Ref. 12) containing

expressions in terms of the amplitude of surface displacement, sound wave

velocities of compressional and transverse mode of propagation, temperature

dependence of heat conductivity and thermal expansion coefficients, elastic

moduli of the material, and the temperature field distribution.

40
The use of lasers to generate ultrasonic pulses at metal surfaces was

first discussed by White (Ref. 13). His analysis considered the expected

temperature rises and the temperature profiles generated in the metal as a

result of laser beam irradiation. Later work (Ref. 14) predicted the form of the

stress pulse likely to be generated once the temperature distribution was

determined (Ref. 15). Two regimes were considered. The first is the

generation of plane elastic waves by thermal expansion at the free and

constrained surfaces due to thermoelastic compressive stresses. The second

regime is the formation of plasma, at very high laser power densities,

resulting in ablation which imparts a momentum impulse to the metal

surface. In general, the heating times are considered to be very short; thus,

thermal diffusion can be neglected. The displacement of the front surface

depends on the velocity of material expansion that produces the sound wave

and the penetration depth of the stress pulse (Ref. 16). For a laser power

density of approximately 1 MW/cm2 and pulse duration of 30 ns, the free

surface displacement is of the order of 1 °A. At higher power densities, of the

order of 100 MW/cm2

, some of the material will be removed by ablation

when the surface temperature reaches the vaporization point. Since the time

to raise the surface temperature to vaporization is relatively long, even for

such high power densities (about 100 ns), the amount of material removed by

ablation is negligible.

Ready (Ref. 17) has shown the temperature distribution obtained with

laser generation of ultrasonic pulses. Solution of the one-dimensional heat

diffusion equation shows that a short high-power density pulse leads to a higher peak surface temperature than a longer, low-power density pulse of

identical total energy.

The amplitude of longitudinal, transverse and Rayleigh surface waves

generated at free metal surfaces by a Q-switched Nd:YAG laser operating at

1.06 pm has been investigated by Aindow et al. (Ref. 18). The acoustic source formed by irradiation of a free metal surface may be approximated as a thin

metal disc. For a source situated deeper in the materials, compressional

waves in all directions would be generated. For constrained surfaces, the

stresses normal to the surface become dominant resulting in an increase in

the generation efficiency of longitudinal waves. When the laser intensity is sufficiently high to cause ablation, the impulse force given to the surface by

the vaporization process enhances the amplitude of the longitudinal wave,

similar to the situation occurring with a constrained surface in the

thermoelastic regime. The Q-switched Nd:YAG laser produced single pulses

with a duration of about 20 ns with energies up to 200 mj. Aindow et al. (Ref.

18) found that a laser pulse energy of 5 mj produced a 2 mV signal from a 33

MHz longitudinal piezoelectric transducer receiver located on the opposite

side of a 1" thick sample of aluminum. For power densities in the

thermoelastic regime, a linear relation was obtained between the laser energy

41
deposition and the receiver transducer output Plasma was formed at the

metal surface when the laser was focused, thereby achieving increased

efficiencies in the longitudinal wave pulse generation. Such an

enhancement is not expected for the generation of transverse waves since

such momentum effects are not a shear generation mechanism. Hutchins et at (Ref. 19) have observed an enhanced efficiency of longitudinal pulse

generation, up to a factor of 40 dB, by application of a constraining overlay.

Surface contamination, such as films and oxide layers, may also enhance the

generation of acoustic pulses. The energy deposited by a Q-switched laser, on

the surface of a metallic sample will be absorbed within the skin depth of the

metal at the irradiating wavelength; however, some thermal diffusion may

occur over the time duration of the laser pulse, thus giving rise to a temperature distribution in a localized small volume. It was shown,

however, that significant temperature rises within the material would not

exist at distances exceeding 10 pm from the surface, even after 100 ns from the

nominal pulse irradiation (Ref. 12,15).

One of the earliest applications of laser-generated ultrasound for flaw

detection was that of Bondarenko et al. (Ref. 20) who used a Q-spoiled ruby

laser emitting single pulses with widths of 30-50 ns and power up to 50 MW.

Calder and Wilcox (Ref. 21) used a single-pulsed, Q-switched ND:glass laser at energies up to 15J with typical pulse duration of about 30 ns. Very strong

ultrasonic echoes were observed but the 15 J laser pulse caused significant

surface damage on polished metallic samples. A focused beam from a ND:YAG laser with a 17 ns pulse duration and an intensity of 40 mj was

applied to perform nondestructive testing on microwelds (Ref. 22).

The risetime of a laser induced elastic wave is comparable with the

risetime of the laser pulse. Availability of such a sharp leading edge suggests

the possibility of high precision measurements of the velocity of propagation

of acoustic waves. Generation of elastic waves, by means of a Q-switched

NdrYAG laser with a 2 mj pulses, has been used to measure the velocity of

extensional waves in metallic glass ribbons subjected to controlled heat

treatments and thus determine the kinetics and mechanism of crystallization

(Ref. 23). The elastic waves propagated along a 200 mm distance were detected

by a 10 MHz quartz transducer.

LASER DETECTION OF ULTRASOUND

Propagation of ultrasonic waves in a medium causes surface

displacements that can be measured optically by exploiting the phase shift of

an optical beam reflected from the surface of the material. When the reflected

beam is appropriately combined with a reference optical beam, from a helium-neon laser, phase changes are converted into amplitude changes that are detectable by a suitable photodiode. These amplitude changes are

42
proportional to the surface displacements on the test sample when an

ultrasonic wave is propagating through it.

Interferometry has been used to measure very small surface

displacements as early as 1900. However, with the advent of lasers, supplying

stable and intense coherent light sources, laser interferometry became an

important analytical tool. The mode of operation of an interferometer, e.g. the Michelson type, is as follows: Light from a coherent monochromatic

source is divided by means of a beam splitter into two secondary beams. One

of these secondary beams travels towards the test sample surface. The other

beam is directed towards a reference mirror located at a fixed distance from

the beam splitter. The two reflected beams are recombined at the beam

splitter, as they return from the reference mirror and test sample,

respectively. The recombined beam is detected by an appropriate detector, e.g., a photodiode. Because the secondary beams originate from the same source,

the laser, sharp interference fringes are produced at the detector. The number

of fringes is a measure of the optical path difference between the two

secondary beams. For small changes, less than a wavelength, the variation in

fringe intensity corresponds to the relative change in position of the surface

of the test sample, since the fringe intensity varies sinusoidally as a function

of the optical path difference. A high sensitivity photomultiplier translates

the changes in intensity to changes in amplitude, or distance.

External noise introduced by mechanical or thermal vibrations may

cause relative changes in the optical path length. However, these noise

signals normally occur at very low frequencies relative to the ultrasonic

frequencies under consideration (1-10 MHz). Apparent path differences may

also arise from relative motion among the optical components of the

interferometer, and even from pressure and temperature fluctuations of the

ambient air. Any specimen motion also shifts the constant path difference

and may displace the fringe pattern, or change the intensity of the detected

recombined beam.

The Fizeau design of an interferometer alleviates part of the difficulties.

To compensate for the random fringe motion so as to maintain maximum

sensitivity the quadrature dual technique is used (Ref. 24,25). The input laser

beam after expansion and collimation is plane polarized and then resolved into

two components by means of a phase retardation plate, X./8. The parallel

component is delayed by 45° with respect to the perpendicular component.

Upon transmission of the returning beams through the retardation plate for a second time, the phase difference becomes 90°, corresponding to circularly

polarized light. Thus, a shifted interference pattern is obtained. In addition,

the square root of the sum of the squares of the high frequency components

(orthogonal to each other) yields the same signal (except for the phase) as would the ideal path difference X/4. Thus, the low frequency random fringe

43
shift is automatically compensated for if both outputs are recorded. Maximum

sensitivity is obtained for any value of the ideal path difference. Such an

interferometer was constructed for determination of changes in elastic

properties and ultrasonic attenuation in rapidly solidified microstructures (Ref.

26, 27). Figure 1 shows a schematic representation of this system.

Stabilization of the interferometer against low frequency mechanical

vibrations can be achieved by an electromechanical feedback arrangement

(Ref. 28, 29). The reference mirror is attached to a piezoelectric device. The

signal from the detector is filtered into high and low frequency components.

The low frequency component is the mechanical noise. The noise signal is applied to a feedback circuit that charges or discharges a capacitor that drives

the piezoelectric mirror mount to eliminate the noise signal. For this type of

apparatus, displacement sensitivities of about 0.5 °A are easily achieved (Ref.

28), which is adequate for most NDE applications involving displacements of

the order of 0.5 - 20 °A.

Compared with other sensors, interferometers offer several advantages.

The sensitive area is definite, and can be made as small as a few microns in

diameter for highly localized detection. The measured quantities are displacement and phase. Independent methods of calibration are applicable.

The bandwidth that determines the fidelity of reproduction of signal

waveshapes is limited not by the physical nature of the transduction process

but by the associated electronic components. In this respect, performance can exceed that of conventional piezoelectric transducers. Small signal resolution

and bandwidth being related, displacements of a few °A are detectable at 20

MHz bandwidth.

DETERMINATION OF DEFECTS IN HOT STEEL BODIES DURING

PROCESSING

It is not clear as yet which type of these laser systems possesses the

features necessary for optimization of a practical system. The desired

intensity of the laser-generated ultrasonic pulse depends to a large extent on

the absorption of the acoustic waves in the test sample. The required

generation intensity will thus affect the basic parameters of the laser detection

device. Among the features to be assessed should be the size and weight of

the laser, ruggedness of construction, power requirements, pulse power and

repetition rate, tube life at operating power level, and cost.

Several state-of-the-art laser beam interferometer detectors of acoustic

waves propagating through metallic solids have been constructed and

employed at Johns Hopkins, the Royal Military College (Canada), and at the

National Research Council of Canada (Montreal). The most recent of these

possesses a frequency bandwidth of 60 MHz and the capability of routinely

detecting waves possessing amplitudes as small as 1 °A. Because of this

44
Fig. 1. Laser-beam generation and dual-laser interferometric detection

of surface and sub-surface ultrasonic waves.

FIBER

SCAN

METHOO FOR DETECTION OF

SHRINKAGE CAVITY AND PIPE IN HOT ROLLED STEEL SLABS

Fig. 2. Non-contact laser generation/ detection ultrasonic system for

inspection of hot steel bodies during processing.
experience, and because of preliminary measurements comparing the

reflectivity of red laser light (helium-neon) with that of blue-green laser light

(argon) from a red hot steel bar, it would be profitable to construct a laser

beam interferometric ultrasonic wave detector incorporating a blue-green

laser. A laser operating in this blue-green wavelength regime is required

since measurements have shown that the signal-to-noise ratio of a red laser

beam interferometers, as used in all prior interferometers, is severely

degraded when used as a detector with a red hot steel body. The reason for

this degradation is that some of the energy radiated by the hot steel matches

the wavelength of the laser itself. The use of a blue-green laser greatly

improves the signal-to-noise ratio under similar circumstances.

In the past, argon lasers have been used exclusively in the blue-green

wavelength range, but recently developed helium-cadmium lasers offer

many advantages. When an argon laser is operated at its maximum power,

tube lifetime is only about 1,500 hours. A helium-cadmium laser operating at

full power is expected to have a tube lifetime of 4,000 hours. The electrical

efficiency of a helium-cadmium laser is five or six times that of an air-cooled

argon laser. Not only do helium-cadmium lasers lower electricity bills, but

more importantly, the cost of a replacement will be significantly greater.

Because argon lasers require higher discharge current (and lower discharge

voltage) than helium-cadmium lasers, they tend to exhibit greater line ripple,

unless specially corrected for. Although argon lasers are usually shorter than

similarly powered helium-cadmium, helium-cadmium lasers are

significantly lighter. These many advantages make helium-cadmium the

laser of practical choice for the proposed research.

For the purpose of identification of the specific laser systems for

generation and detection of ultrasonic waves in hot steel slabs, the effect of

nonadherent oxide layers on the steel slabs should be evaluated. It is not at

all obvious that an oxide layer (scale) on the surface of the steel slab will

interfere with the ultrasonic measurement. The mechanical and physical

properties of the oxide layer, e.g. temperature, ductility, thickness, adherence

to the surface, etc., determine whether the presence of an oxide layer is beneficial or not. In certain cases, where a compatible laser irradiation is generated, such a layer may act as an artificial overlay on the irradiated

surface to produce a shock wave in the test sample. Under these

circumstances, the generated ultrasonic pulse will be highly enhanced.

A series of experiments, properly designed and executed, should

provide the necessary information for the evaluation of the noncontact

methodology. This methodology has a great potential that should be

exploited, particularly for solving the important problem that has eluded any

realistic solution so far, namely, the inspection and characterization of hot

steel bodies during actual processing.

46
Noncontact ultrasonics, employing laser generation and laser

interferometric detection of sound waves can be remotely operated by means

of optical fibers. Thus, the delicate instrumentation (lasers, etc.) can be located

several hundred yards away from the steel processing halls. Furthermore,

ultrasonic signals can easily be digitized and fed back into monitoring and

control units of the metal processing systems. Therefore, a high degree of

automation can be achieved while improving the reliability, quality assurance and profitability of the hot processing of steel. Figure 2 shows a schematic presentation of such a noncontact laser generation/ detection

ultrasonic system for inspection of hot steel bodies during processing.

47

REFERENCES

1. H. E. McGannon, "The Making, Shaping, and Treating of Steel". U.S. Steel

Corporation, Ninth Edition, Herbien and Held, Pittsburgh, PA. (1971).

2.

"Research Needs for Process-Control Sensors", American Iron and Steel

Institute General Research Committee—Task Group on Steel-Industry

Research Priorities for Process Control and Sensor Development (January

1981).

3. B.E. Droney and T. J. Pfeiffer, Materials Evaluation. 38. 1130 (1981).

4. R. C. Booth and R. N. Cressman, Materials Evaluation, 39, 1130 (1981).

5. E. P. Papadakis, L. C. Lynnworth, K. A. Fowler and E. H. Carnevale,

T. Acoust. Soc. Am. 52. 850 (1972).

6. K. R. Whittington, British T. NDT, 242 (1978).

7. K. Miyagawa et al.,

"Nondestructive Testing Systems for Pipe and Tube

Mills at Nippon Corp.". Nippon Steel Techn. Report No. 14 (1979).

8. J. M. Mandula, Materials Evaluation, 30, 49 (1972).

9. K. G. Bergstrand and P. Nilsson, Iron and Steel Engineer, 57, 1 (1980).

10. J. Pachtem and J. Cornet, "Resultats Pratiques d'une Gammagraphie en

Exploitation Industrielle". Commission des Ingenieurs de Laminoirs,

(Feb. 1968).

11. J. C. Albert, J. Dumont-Fillon and J. Pinaud, Proc. 8th World Conf. on

NDT, Cannes, France, (Sept. 1976).

12. J. F. Ready, T. Appl. Phvs. 36. 462 (1965).

13. R. M. White, T. Appl. Phvs . 34, 3559 (1963).

14. L. S. Gournay, T. Acoust. Soc. Am. 40, 1322 (1966).

15. A. M. Aindow, R. J. Dewhurst, D. A. Hutchins, and S. B. Palmer, in 1980

European Conf. on Optical Systems and Applications, Utrecht, Holland

1980, European Conf. Abst. Vol. 41, Paper N6.

16. J. C. Bushnell and D. J. McCloskey, T. Appl. Phvs. 39, 5541 (1968).

49
17. J. F. Ready, "Effects of High Power Laser Radiation", Academic Press, New

York (1971).

18. A. M. Aindow, R. J. Dewhurst, D. A. Hutchins and S. B. Palmer, T. Acoust.

Soc. Am. 69, 449 (1981).

19. D. A. Hutchins, R. J. Dewhurst, S. B. Palmer, Ultrasonics. 19. 655 (1976).

20. A. N. Bondarenko, Y. B. Drobot and S. V. Kruglov, T. NDT. 12, 655 (1976).

21. C. A. Calder and W. W. Wilcox, Mater. Evaluation, 38. 86 (1980).

22. Y. Bar-Cohen, T. NDT. 21, 76 (1979).

23. M. Rosen, H. N. G. Wadley and R. Mehrabian, Scripta Met. 15, 1231 (1981).

24. E. R. Peck and S. W. Obetz, T. Opt. Soc. Am. 43, 505 (1953).

25. D. Vilkomerson, Appl. Phvs. Lett. 29. 183 (1976).

26. M. Rosen, C. H. Palmer and S. Fick (private communication).

27. M. Rosen and E. Horowitz, The Johns Hopkins University, Center for

Materials Research Report CMR-NDE 7 (Sept. 1982).

28. C. H. Palmer and R. E. Green, Jr., Appl. Opt. 16, 2333 (1977).

29. C. H. Palmer and R. E. Green, Jr., in

"Nondestructive Evaluation of

Materials",

J. J. Burke and V. Weiss, editors, pp. 347-378, Plenum Press,

NY (1979).

50
Impedance of a Coil in the Vicinity of a Crack

Arnold H. Kahn

National Bureau of Standards, Washington, DC 20234

Calculations are presented for the impedance of a coil as it is moved in the vicinity of a

v-groove crack in the surface of a metallic slab. The coil is modeled as a pair of parallel

wires, oriented parallel to the crack, carrying equal and opposite currents. The

inhomogeneous electromagnetic fields in the air above the slab and in the metal are determined by the boundary integral equation (BIE) method. This approach leads to a

pair of coupled integral equations for the tangential components of the electric and

magnetic field vectors on the surface of the slab containing the crack. The solutions,

which are obtained by standard method of discretization, are valid for arbitrary ratio of

crack or coil dimensions to skin depth. Illustrations are presented of the Poynting vector

distribution over the surface of the metal, including the crack faces. A plot of the complex

impedance is given in the form of a coil scan across the crack.

1. Introduction

In the design of electromagnetic NDE

systems for the detection and examination

of cracks or other defects in conducting

materials, it is necessary to have a

quantitative description of the electric and

magnetic fields in the vicinity of the defect.

In practice, the fields are produced by an

exciting coil, the impedance of which is used to provide the detection signal. (The

voltage induced in a secondary pickup coil

may also be used.) In previous work by the

author [1,2], the fields in the vicinity of a crack were calculated for models based on

excitation by a spatially uniform applied ac magnetic field such as would be found in

the interior of a solenoid. The present work

offers an improved description of the fields

through the introduction of nonuniformity

of the applied field due to finite coil size and

the inclusion of coil position relative to the

crack.

Recently there has been significant

activity in the development of theoretical

modeling in electromagnetic NDE. The

finite element method has been applied by

Ida and Lord [3] to the cylindrical

geometry of reactor tubing. Studies have

been presented by Auld et al., Kincaid et al.,

Bahr, and other [4] on experimental and

theoretical considerations of crack detection

and coil design. A principal difficulty is the

calculation of signals when the

electromagnetic skin depth and crack size are of comparable magnitude, which is the

domain of greatest sensitivity. The twodimensional model of this paper represents a contribution toward the solution of this

problem. A full three-dimensional

treatment may be possible as new

computing capabilities are developed [3].

2. Description of the Model and

Theoretical Formulation

The calculations described in this paper

are based on the following model:

We consider a flat surface with an

infinitely long, symmetrical v-groove

representing a surface crack, the material

is homogeneous and uniform in

conductivity. A pair of wires carrying

equal and opposite currents is located above

the slab and is oriented parallel to the

crack. The wires are infinitesimal in

thickness and infinite in extent. This

51
simplified model of an eddy-current testing configuration allows a two-dimensional calculation of the

impedance signals due to the crack. The calculation

will allow for the effect of crack dimensions; coil dimensions, elevation, and displacement; and the

material parameters of the metal. This is an improvement over calculations in which the exciting field is spatially uniform.

By solving for the electromagnetic fields first on the

surface of the metal and then at the exciting wires, we

obtain the impedance due to the presence of the

metallic region. If the problem is solved for a plane

surface without the crack, then the additional

impedance due to the crack may be obtained by

subtraction. Also, by solving for different positions of

the wires representing the detection coils, we may

obtain the impedance signal on traversing the crack

and also the signal due to liftoff effects. The model is illustrated in figure 1. The circles above the surface represent the wires, with + and -

indicating the direction of the impressed current I0

e'‘u

\

where &> is the angular frequency and t the time. The

current is held at fixed amplitude /o> according to the

usual procedure for eddy-current NDE. In the figure

additional parameters are shown: A is the separation

between the wires, H is the height of the coil above

the plane, P is the center position of the coil relative to the crack, D is the depth of the crack, and F is the

half-width of the crack opening.

Because of the symmetry of this two-dimensional

model, the electric and magnetic fields may be derived

from a vector potential, A, which has only one

component, A z [5], where the z-direction is parallel to the wires and the crack. If the wires were not parallel to the crack, a full three-dimensional analysis would be

necessary. The vector potential is thus of the form

A(x,y)e~

M', where A is complex to represent phase

relations with respect to the exciting current.

In the region above the conductor the vector

potential satisfies a Helmholtz equation. However, at the frequencies of eddy current testing the transit time

for wave propagation across the region of the crack is negligibly small and a quasi-static approximation is satisfactory. Thus, in the region above the metal slab the vector potential satisfies the Laplace equation,

V

2A=0, (1)

except for the singularities at the wires. Below the

surface, in the metallic region, the Helmholtz equation

is obeyed,

(V2 q- k

2)A =0, (2)

where

k

2 =icro)pi (3)

is the square of the propagation constant, cr is the

electrical conductivity, and p is the magnetic

permeability. Here too, displacement currents are neglected, in this case because the ohmic currents represented by the k

2 term are so much larger. At the boundary surface, including the faces of the crack, the

© © ~

h

A ~H h

Figure 1-Configuration of model

and parameters for the calculation of the impedance signal due to a crack.

ll

D Crack Depth A Wire Separation

2F Crack Opening H Wire Elevation P Wire Position Relative

to Crack

52
usual conditions of continuity of tangential E and H,

and normal B and D must hold. In terms of the vector

potential, these conditions are equivalent to the continuity of A and \/\i dA/dn, where dA/dn is the normal derivative of A at the interface. To summarize,

the vector potential must satisfy the Laplace equation

in the upper region, the Helmholtz in the lower

region, and conditions of continuity at the interface,

and it must approach the known form of the impressed

field in the vicinity of the source wires.

The method of solution selected in this paper is an

extension of the boundary integral equation (BIE)

approach [6,7]. This method, usually applied to a

single region, has been used by the author [2] in eddycurrent problems involving excitation by a uniform ac

field. In the present application the method leads to a

pair of coupled Fredholm integral equations of the

first kind, as follows: By application of Green’s theorem we express the vector potential in the upper region in terms of the source fields and the values of A and its normal

derivative dA/dn on the bounding curve:

A(r) =y(r)

- J

C(rS) £.

(S1 dS\ (4)

in which dS' is an element of arc in a planar cross- section normal to the surface of the metal. In the above, is the vector potential due to the source

wires as if the metallic region were absent. The two

integrals give the change due to the induced currents below the boundary. They are taken over the

boundary n is the unit normal vector pointing out of

the upper region. The remaining boundary closure at infinity makes no contribution since the fields decay

with sufficient rapidity. Green’s function for the

Laplace operator is given by

G(r,r') = - 1/2-tt log | r-r' |; (5)

it satisfies V

2 G(r,r') = -S(r-r'), (6) where 6 is the two-dimensional Dirac delta function.

For the two-wire case treated in this paper, the source

field 5 has the form

S(r)=/0[G(r,r+

)-G(r,r)],

where r + and r_ are the positions of the wires which

carry the exciting current, 70, parallel and anti-parallel to the z-direction, respectively. In the metallic region

A(r)=- f

(r,S

'> A(S')dS'

J dn

+ J

^(r,S')^f

S,) dS\ (7)

where & is now the two-dimensional Helmholtz

Green’s function,

&(r,r) —(i/4-)HQ°\k \ r-r | ), (8)

where H0

{]) is the Hankel function of the first kind,

order zero. It, too, has a logarithmic singularity and

satisfies the Helmholtz equation with a source,

(V2 + k

2) 6?(r/) = -8(r-r'). (9)

This latter Green’s function contains the complex k

and represents a damped outgoing cylindrical wave at large values of r-r'. In eq (7) we have retained the same direction of the normal vector n; hence the unusual sign convention on the right hand side. The BIE method prescribes letting r approach the surface to obtain the fields on the bounding surface. When we let r=S, a well-defined expression is obtained if we use the Cauchy principal values for the singular integrals and replace A on the boundary by

A/2. For nonmagnetic materials A and dA/dn are both

continuous across the boundary, and we shall so

restrict the present calculations. The resulting BIE’s

are

hA(S)-j ^,

(S’ S ) A(S')dS'

+ J

G(S,S') dS' = y(S) (10)

M(S) + J

A(S')dS'

- j

& (S,S') ) dS' =0 (11)

This is a coupled pair of equations for unknowns T(S)

and dA(S)/dn on the interface. We may look on the

53
inhomogeneous term y(S) as the driving force for the system. When A and dA/dn are found on the

boundary, then the field A may be constructed at any

point above the interface by application of eq (4), or below the interface by application of eq (7). The ultimate objective is the determination of the

impedance per unit length induced in the wires by the

presence of the metallic region. The time-average of

the power per unit length delivered by the exciting

wires is given by the complex expression

P=\ f E'-J*da,

where J is the (constant) current density in the wires, E' is the electric field at the wires produced by the induced currents, and da is an element of cross- sectional area normal to the wires. E is derived from

the vector potential A

of the induced currents,

a '=A-y,

by the usual relation

E' = icoA'.

Hence, for a set of idealized line-wires, denumerated

by the index z, we have

P= \1IfitaA\.

Connection with conventional circuit parameters can

be made by expressing P in terms of the currents,

voltages per unit length, and impedance per unit length of the wires. Under the constant current assumption of eddy current testing, we have

p=Yivi if=k'iiffzt,

where Vi is the voltage per unit length induced in the

zth wire and Z, is the extra impedance in the zth wire

due to induced currents in the metallic region. Finally

we obtain

Z, =zco/4'//,

for the impedance per unit length in the zth wire,

caused by the induction. Now Ax is evaluated at the zth wire and can be computed by use of Green’s theorem

after A and dA/dn have been found on the interface. Thus we have a method of computing the extra impedance seen by each wire due to the presence of

the metal below. These impedances may now be

calculated with and without a crack being present.

3„ Numerical Treatment

The coupled integral equations are solved by an

application of the method of moments [8]. The

solution is expressed as a linear combination of a finite set of basis functions with unknown coefficients. The

coefficients are determined by requiring that the

integral equations be satisfied at a number of points

equal to the number of unknown coefficients; i.e., point-matching is used. For the basis functions, the elements F shown in figure 2 were used, after the method of Harrington [8].

/

/

/

A

' \

/ \ \\

\

/

/ 4

/ /

/

\

\ \

\

\ v

Sn-i Sn+ i

Figure 2-Triple-pulse hat function F used in the numerical calculations. The illustration shows the function F(S-SJ. The dashed

line is the common triangular hat-function.

We approximate the solution for the vector potential

and its normal derivative on the interface by the finite sums:

A(S) = IA,F(S-S,) (12)

and

dA(S)

dn = 2N,F(S-Sl) (13)

These expansions are introduced into the integral

equations, eqs (10) and (11). The integration over each

element is carried out by use of the midpoint rule for the entire integrand of each flat section of the fundamental element, except when the Green’s

function is singular, i.e., when S—Sj.

When a singular integrand occurs in the evaluation

of eq (10), the integration of the logarithmic Green’s

function is performed exactly. When a singular

integrand occurs in eq (11), the dominant logarithmic

part of the Hankel function is used for the evaluation.

The width of the elements is not restricted. It was

found practical to use a fine grid where the solution

54
was large or varying rapidly, and a coarse grid

elsewhere. With these approximations and the pointmatching, the coupled integral equations are reduced

to a linear algebraic system of the form:

when multiplied by the Hankel functions of the integrand. The solution of the linear equations was

obtained by Gaussian elimination without pivoting.

The logarithmic singularities of the Green’s functions

associated with the diagonal elements of the matrices

allow this economical simplification. The solutions were considered to have converged when further refinement produced an insignificant change in the physical results, usually about 1 percent. Typically the dimension of the square matrices ranged from 200 to 300.

4. Coil Impedance in the Absence

of a Crack

In these equations, each doubly-subscripted term

corresponds to that part of the integrations of eqs (10)

and (1 1) connecting element j and matching-point i. The calculations were first attempted with square

pulse functions as the basis set. It was found that the

solutions were unstable in the vicinity of the crack

corners and near the location of a grid-size change.

The use of the triple pulse element is equivalent to doubling the number of points in a pulse function

calculation, but applying the constraint that the

solution at each point be averaged with its two nearest neighbors. In addition to reducing the dimensions of

the needed matrices, this has a smoothing effect and

leads to solutions which are stable as the grid size is decreased. The triple pulse basis function may be

looked upon as an approximation to the common

triangular hat-function, shown in figure 2 by the

dashed line. The hat-function yields a piecewise

trapezoidal approximation to the solution which

would be superior to the present form, but its application is precluded because of nonintegrability

The radiation field of an oscillating dipole above a conducting earth was a problem first attacked

successfully by Sommerfeld [9]. Analytical solutions

have been given for finite coils by Dodd et ai [10,11].

These solutions are in the form of integrals over

various Bessel function arguments. Numerical

evaluation is possible; analytic evaluation is in terms of

asymptotic series. The same methods can be applied to

this problem of a pair of parallel wires over a plane.

However, the approach of this publication is readily

applicable in the absence of a crack. Solving an

integral equation requires a greater computing effort than evaluating an integral solution for the lesser problem. However, it is quite useful to have the programs available as a byproduct of the crack case. In this section we examine the results of calculations

for the parallel wire coil above a flat conducting half- space, calculated by the boundary integral equation

method.

In figure 3 we show the results of a typical

calculation. For this case, and all others reported here,

Figure 3-Poynting vector on the surface of a metallic slab in the absence of a crack. The

coordinate x is along the flat surface of the metal. Distances

are in units of the skin-depth and the Poynting vector is in units of 1CT

3

. The exciting wires are located at ±0.5 6 and

are at an elevation of 0.5 8.

PLOT SURFACE

O o

55
lengths are in units of the skin depth 8, where

8 — V2/cra>ju0, (16)

and the symbols under the radical are the same as before. The complex Poynting vector S, represents the time average of the complex energy flux, in our

application, across the surface of the conductor. In the

units we are using, we have

S = hExH*

i_ JXqOJ 2 8

where A and dA/dn' are calculated by solving the

coupled boundary integral equations. While our

principal interest is in the impedance change of the

exciting wires, the Poynting vector plots show a detailed picture of the radiation field. The plots are

useful for assessing the convergence of the

calculations as well as for showing the regions of the

test material where the significant absorption and field penetration take place.

5, Coil Impedance With a Crack

The presence of the crack adds two more

parameters to the required inputs to the calculation.

We treat only a symmetric v-groove crack which we

specify by its depth and the half-width at its mouth.

The calculations are performed in the same way as without the crack, with the only difference being that

the needed matrices are larger in dimension and

somewhat more complex in preparation. The

algorithms for the solution are identical to those of the previous case. The output of the program is the impedance per unit length of the wires, with the crack

present. In addition we may inspect the complex

Poynting vector on the surface of the crack as well as on the flat surface of the test material. For the initial investigations we selected a crack

depth of 2,0 8 and an opening of half-width 0.25 8. The coil wires were taken as having a separation of

1.0 8 and at an elevation of 1.0 8 above the plane.

These dimensions correspond to the physical situation of a No. 30 AWG wire insulated pair in close contact,

elevated one radius above the contact with the plane,

and driven at a frequency of 1 10 kHz. The relevant

parameters for this model applied to aluminum are given in table 1.

Table 1. Parameters for model calculation based on aluminum

at 1 10 kHz.

Resistivity P 2.82 X10'8 n-m(20‘C)

Conductivity <r(=l/p) 3.54X107 ft-W

Skin depth 8 0.255 mm

Crack depth ( = 28) 0.51 mm

Crack half-opening (8/4) 0.064 mm

Wire radius (8/2) 0.13 mm

The calculations were performed for a range of values

of the parameter P, the location of the coil center

relative to the crack. Figures 4, 5, and 6 show the Poynting vector for the values P=2.5 8, 0.5 8, and 0.0

Io dA*

dn'

(17)

FLAT SURFACE CRACK FACES

Figure 4-Poynting vector on the surface of a metallic slab with a crack The coordinate S is along the faces of the crack, which is shown as folded open in the right-hand figure. The shaded band indicates the location of the v-groove crack. The lateral distance between the coil and the crack, P, is 2.0 8 and the half-opening, F, is 0.25 5. All other parameters are as in figure 3.

56
PLAT SURFACE CRACK FACES

Figure 5-Poynting vector on the surface of a metallic slab with a crack. The lateral distance between the coil and the crack, P, has the value of 0.80 S; all other parameters are as in figure 4.

FLAT SURFACE CRACK FACES

Figure 6-Poynting vector on the surface of a metallic slab with a crack. The lateral distance between the coil and the crack, P, has the value

0.0 S; all other parameters are as in figure 4.

S respectively, illustrating the deformation of the

fields as the coil is brought up to the crack. These

figures correspond to the same wire separation and

elevation as in figure 3, the case with no crack.

Qualitative examination of the figures shows that in the presence of a crack, a portion of the integrated

Poynting flux is “stolen” from the nearer of the peaks

in the field distribution. The Poynting flux at the corners is somewhat increased over the value that would occur at that position if no crack were present.

Inside the crack, the Poynting flux decays to zero in approximately one skin-depth. This is quite the opposite behavior to that which occurs in the case of a uniform //-field parallel to the crack [2]. In that case, the Poynting vector is greatest at the tip and vanishes

at the comers. From a series of calculations like these, the coil impedance per unit length was obtained for numerous

positions of the coil. The phase and magnitude of the impedance are shown in the plots of figure 7, in the

57
Phase of Coil Impedance

_ Magnitude of Coil Impedance

CNJ

Figure 7-Plots of the phase and magnitude of the crack impedance

signal as a function of the lateral displacement of the coil center

relative to the crack. Parameters are as in figure 4.

form of a scan across the surface of the slab. These

curves would be extended symmetrically for negative

values of the coil displacement. The asymmetric signal

obtained from an opposed pair of coils [12] can be

obtained from these curves by computing a differential scan corresponding to the coil pair separation.

The author wishes to express his thanks to Dr. B. Auld and Dr. C. Fortunko for most valuable

discussions concerning crack detection. He is also grateful to Dr. S. Gershovits for interesting

suggestions.

References

[1] Kahn, A. H.; Spal, R.; Feldman, A. J. Appl. Phys. 48, 4454 (1977).

[2] Kahn, A. H. in Review of Progress in Quantitative Nondestructive Evaluation 1, D. O. Thompson and D. E. Chimenti, eds. ( Plenum Press, New York, NY, p. 369

[3] Ida, N.; Lord, W. IEEE Computer Graphics and Appl. 3, 21 (1983).

[4] Review of Progress in Quantitative Nondestructive Evaluation 2, D O. Thompson and D. E. Chimenti, eds., (1983).

[5] Landau, L.. D.; Lifshitz, E. M. Electrodynamics of Continuous Media

,

Pergamon Press, New York, NY (1960), p. 121.

[6] Rizzo, F J. in The Boundary-Integral Equation Method:

Computational Applications in Applied Mechanics, T. A. Cruse and F. J. Rizzo, eds., American Society of Mechanical Engineers, 1975, p. 1.

[7] Jaswon, M A.; Symm, G. T. Integral Equation Methods in Potential Theory and Electrostatics,, Academic Press, London (1977).

[8] Harrington, R. F. Field Computation by Moment Methods, Macmillan, New York, NY (1968).

[9] Sommerfeld, A. Partial Differential Equations in Physics, Academic Press, New York, NY (1949), p. 236 et seq. [10] Dodd, C. V.; Deeds, W. E. J. Appl. Phys. 39, 2829 (1968).

[11] Dodd, C. V.; Deeds, W. E.; Luquire, J. W. Int. J. Nondestructive Testing 1, 29 (1969),

[12] Fortunko, C. M.; Padget, S. A. Technical Activities 1982, Office of Nondestructive Evaluation, NBSIR 82-2617, p. 57 et seq.

58
3. WORKSHOP SUMMARIES

Three concurrent discussion groups (A, B, and C) were formed for

consideration of the internal discontinuity problems that were most

frequently encountered in the following hot areas:

Group A: Solidification

Group B: Intermediate Forming

Group C: Finishing

Discussion questions were provided for each group session. Each were

discussed and conclusions were drawn concerning the priority of detection

and the best sensor technology to explore.

3.1 Solidification - Chairman, J.R. Cook, ARMCO Inc.,

Recorder, H.N.G. Wadley, NBS

Question 1. What are the internal discontinuities of importance during

continuous casting? Can they be prioritized on the basis of yield, controllability,

feed forward /feedback control, quality?

• The general categories of internal discontinuities in prioritized order of

importance during continuous casting are:

1. Inclusions

2. Cracks

3. Porosity (center looseness)

4. Center line segregation

• Calcium aluminum silicate inclusions in the 10 pm - 2 mm size

range are the highest detection priority. Their occurrence is indicative of a breakdown in shroud gases just prior to casting.

Smaller (<10 pm) aluminum oxide inclusions from aluminum

killing are a less important detection priority. In either case, the

appearance of significant inclusion concentrations flags a need for

corrective action further back in the process and a region of

unsound steel for future cropping. Corrective actions include

adjusting shroud gas composition, ensuring correct degree of

electromagnetic stirring, control of vortexing and avoidance of

nozzle wear.

59
• Due to their difficulty of hot detection with todays technology,

surface and subsurface cracks are, a major source of customer

claims. They are responsible for extensive rework and

downgrading of substandard product during downstream

processing. Principal types of cracks encountered are:

1. Starcracks: caused by a local excess of copper (pickup from mold).

The cracks are starshaped, emanating along grain boundaries

outward from the region of enhanced copper concentration.

2. Oscillator cracks: surface cracks nucleated at the base of mold

oscillator marks due to a local stress concentration. They are usually a manifestation of improper casting speed /cooling practice.

They can also be caused by misalignment and bending.

3. Surface Rip/Tears: infrequently encountered, usually due to poor

lubricity in the mold.

4. Longitudinal cracks: usually caused by poor heat transfer due to localized overcooling and sometimes misalignment.

5. Subsurface cracks: longitudinal cracks, radial cracks and core cracks

that may refill with steel forming potential internal discontinuities.

• Center looseness and shrinkage porosity may arise from an improper

ratio of steel quantity to casting speed combined with the wrong

cooling practice. Depending upon the severity, the condition

sometimes can be corrected by subsequent reduction of cross

sectional area.

• Center line segregation of nonmetallics may occur because of

vortexing in conjunction with cooling of the liquid center. Usually

electromagnetic stirring avoids the problem. If it is detected, it is a manifestation of an equipment malfunction, improper control of

withdrawal speed and/or incorrect metallurgical length.

Question 2. Where do internal discontinuities occur during continuous

casting? Can optimum sensor locations be identified for each type of

discontinuity?

• The most deleterious inclusions form in liquid steel prior to its solidification. Teeming and other operations are used to reduce

their concentrations to acceptable levels prior to entering into the

mold. Mold fluxes then take over the role of entrapping and

removing inclusions prior to solidification. Sensors located both to

60
monitor the efficiency of teeming and mold fluxing are required.

The former would be located in the tundish, the latter near the

mold exit.

• Cracks form in steel while it is in the mold, or during secondary

cooling prior to cropping. Sensors ideally should operate just before

and just after secondary cooling so that the origin of cracking

problems may be identified.

• Center looseness/porosity occur at the freeze-off point below the

mold. A sensor to detect this phenomenon should be located at this

point, provided it is known.

Question 3. Do internal discontinuities occur in higher concentrations

during transient conditions (whether random disturbances or changes due to control action)? What transients are of importance? What discontinuities do

they produce? Do we need sensors only when we change process conditions?

• There are three primary transient conditions that can cause internal

discontinuities to occur in higher concentrations:

1. At the beginning of the cast, the material behind the starter bar is unsound and must be cut off and discarded. The precise point of

the transition of unsound material, containing a high incidence of

impurities, is at present estimated (not measured).

2. At any time, there is an equipment failure such as nozzle erosion

spray nozzle stoppages or loss of stirring effectiveness. *

3. During the intermix zone during grade changes. In some cases, the

process must be restabilized to a new operating point.

• The transient conditions above can impact internal cleanliness (with

the introduction of both large and small inclusions with center line

segregation), internal cracking and/or center looseness (porosity).

Since most upsets occur at random, it follows that a continuous

scanning sensor is preferred. However, it is recognized that the size,

orientation and types of internal discontinuities may, for some

conditions, be predictable. For example, the starter bar and the

intermix zones are known events. In each case, the types of internal

discontinuities are known; therefore, the sensitivity requirements of

sensing device may be different for these situations. Perhaps a study

would be advantageous to rank the cost effectiveness of each sensing

situation, and address collaborative program objectives on this as a

priority.

61
Question 4. What benefits accrue from online detection of internal

discontinuities? Can we quantify these benefits (% yield, % productivity,

added product value)? How much should a sensor cost?

• Presently, no sensing/ feedback control exists for production casters.

Statistical process control (SPC) or quality management programs

(QMP) are used to establish an operational "window" that results in

the highest quality. Transients result in temporary excursions out of

the window. Faults may also result in changes to the cooling system,

flux efficiency, etc., that also take the caster outside its intended

operational window. It may take days to weeks to recognize this.

• Internal discontinuity sensors could provide real-time indication of

harmful excursion outside operational windows. Sensors with

sufficient resolution /correct positioning could be a corner stone for

narrowing the "window" and thus reducing caster operation costs,

and improving the overall quality of material cast within window

conditions.

• Considerable benefits are to be anticipated from the implementation

of internal discontinuity sensors. These vary from caster to caster

but nationwide could result in a 2% yield improvement.

• A sensor should "pay back" within six months to a year of

implementation if it is to be economically justifiable in todays

climate. Thus it must cost at least less than the value of a 1% yield

improvement and preferably a factor of 3-5 less than this. Given the

industry-wide need for such sensors, and their probable applicability

to aluminum and other metal casters, a collaborative research

program is probably called for in order to spread out research and

development costs.

Question 5. what is the optimum sensor methodology for each class of

internal discontinuity?

• Several sensor methodologies show promise. They include:

1. Ultrasound

2. Eddy currents

3. X-ray tomography

• To achieve the spatial resolution needed in thick steel, x-ray

tomography, would require very intense sources and numerous

62
detectors. Such a system could cost several millions of dollars

unless emerging NDT systems can, in some way, be modified. Both

ultrasonic and eddy current systems could be configured within the

allowable cost range.

• Ultrasonic approaches are probably the best for center looseness,

center line segregation, inclusions and some cracks. Traditional

pulse-echo approaches fail for surface cracks but pitch-catch or

surface wave techniques may be satisfactory. Eddy current methods

show potential for surface crack problems but no other.

3.2 Intermediate Forming, Chairman, C.D. Rogers, U.S.S. Corporation,

Recorder, M. Linzer, NBS.

Question 1. What are the internal discontinuities of importance during

intermediate forming? Can they be prioritized on the basis of yield,

controllability, etc? Are classes of internal discontinuity product specific?

• Three classes of defect should be detected. In order of priority they

are:

1. Subsurface inclusions greater than 150 (im located within 12.5 mm

of the surface. These inclusions are mostly nonmetallics introduced

in the molten metal prior to entry into the caster.

2. Surface cracks greater than 12.5 mm depth and 150 mm long (after

rolling).

3. Subsurface inclusions and porosity with dimensions greater than 2.5

mm at depths greater than 150 mm. These defects can occur in

almost any product. The surface defects can result in poor surface

quality of strip product. Internal discontinuities can be particularly severe for defense specification grades and product destined for

drawing/forming.

Question 2. What sensor methodologies are best suited to online detection of

each class of internal discontinuity?

• Ultrasonic techniques show the best promise for detecting subsurface

inclusions, cracks and porosity.

• Eddy current methods may prove useful for detecting surface cracks,

though ultrasonics may be the preferred approach ultimately.

63
Question 3. What are the benefits accruing from online detection of internal

discontinuities? Can these benefits be quantified (% yield, % productivity,

added product value)? How much should a sensor cost?

• Optimized cropping/ product diversion would maximize

productivity while maintaining acceptable quality.

• 100% sensor inspection would provide valuable information about

casting practices and could be fed back to optimize continuous casting

practices.

• Optimized reduction to seal center looseness and avoid downstream

quality problems during piercing, drawing and/or forming.

3.3 Finishing. Chairman, W. Wilson, Weirton Steel Corporation, Recorder,

R.B. Clough, N.B.S.

Question 1. What are the internal discontinuities of importance during

finishing? Can they be prioritized on the basis of yield, controllability, etc?

Are classes of discontinuity product specific?

• The internal discontinuities of importance during finishing are

inclusions, piping and center line cracking. These can be prioritized

on the basis of yield, detectability, cost, controllability and the stage of

manufacture. Because of all of these factors, the classes are usually

product specific.

Question 2. What can be done when discontinuities are detected?

• The product can be accepted if it meets certain specifications which

are a function of vendor and market conditions as well as production costs or rejected. It is desirable to reject such products as early in the production cycle as possible to reduce production costs. The level at which the product is rejected is critical and will depend

on the technology of detection and economics of the specific

situation. Early detection and remedial actions to upstream

processes is the preferred strategy for quality improvement.

Question 3. What sensor methodologies are best suited to online detection of

each class of internal discontinuity?

• Acoustic methods are best suited for thick steel sections. These

methods are also more sensitive to internal discontinuities than

radiographic methods, which do not provide good contrast for cracks

64
that are closed. Eddy current methods are limited to surface

discontinuities.

Question 4. What benefits accrue from internal discontinuity detection

during finishing operations? Can these benefits be quantified (% yield,

% productivity, added product value)? How much should a sensor cost?

• The benefits of internal discontinuity detection at this stage of

processing are not economic, at least in the short term, due to higher

recycling costs and decreased productivity. Benefits accrue in the

long term however, due to decreased customer dissatisfaction and

enhanced quality.

• It is desirable to detect internal discontinuities prior to finishing

operations to reduce production costs.

• A sensor ideally should not cost over $1,000.

Question 5. Do other quantities (e.g., grain size) need to be measured during

finishing?

• The consensus of the group was that these would not provide

sufficiently useful information to justify costs. Unless such

information were required for specification purposes, such

measurements do not appear necessary.

65
*

.
NBS-114A (rev. 2 »ac) U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See instructions)

1. PUBLICATION OR

REPORT NO.

NBSIR 88-3731

2. Performing Organ. Report No. 3. Publication Date JUNE 1988

4. TITLE AND SUBTITLE

Report of Workshop - Internal Discontinuity Sensor Needs for Steel

5. AUTHOR(S)

H. N. G. Wadley

6. PERFORMING ORGANIZATION (If joint or other than NBS, see instructions)

NATIONAL BUREAU OF STANDARDS

U.S. DEPARTMENT OF COMMERCE

GAITHERSBURG, MD 20899

7. Con tract/ Grant No.

8 . Type of Report & Period Covered

9.

SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City. State, ZIP)

The American Iron and Steel Institute

1133 15th St. N.W.

Washington D. C. 20005

10.

SUPPLEMENTARY NOTES

| | Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11.

ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bi bliography or literature survey, mention it here)

Operators, engineers and researchers representing American Iron and Steel

Institute member companies met with scientists and engineers from federal

laboratories, universities and sensor vendor companies to determine the formation

mechanisms of internal discontinuities in continuous cast steel, identify sensors

for their detection and control strategies for their elimination.

The development of online sensors to detect inclusions and cracks within the

caster was determined to be urgently needed for real time quality control.

Outputs from such sensors could be used to detect inert gas shroud cracking

(the most common source of the most deleterious inclusions) and cracking problems

and facilitate feedback control. Of all the potential sensor methodologies

addressed, ultrasonic approaches similar to those developed for porosity work and

also related to those under development for internal temperature/solidification

interface measurement were considered the most promising. Collaborative

programs need to be instituted to prove the ultrasonic approach and develop data

for the design of a prototype sensor.

12. KEY WORDS (Six to twelve entries; alphabetical order; capital'/ ze only proper names; and separate key words by semicolon s) continuous casting; eddy currents; inclusions; porosity; solidification;

ultrasonics.

13. AVAILABILITY 14. NO. OF

PRINTED PAGES

fxl Unlimited

70

| For Official Distribution. Do Not Release to NTIS

| | Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C.

20402. 15. Price

[_Xj Order From National Technical Information Service (NTIS), Springfield, VA. 22161 $13.95

USCOMM-DC S043-P80