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U.S. DEPARTMENT OF COMMERCE / National Bureau of Standards

Technical Digest

Symposium on Optical Fiber

Measurements, 1986

Sponsored by the National Bureau of Standards

in cooperation with the IEEE Optical Communications Committee

and the Optical Society of America
Tm he National Bureau of Standards' was established by an act of Congress on March 3, 1901. The

Bureau's overall goal is to strengthen and advance the nation's science and technology and facilitate their effective application for public benefit. To this end, the Bureau conducts research and provides: (1) a

basis for the nation's physical measurement system, (2) scientific and technological services for industry and

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The Bureau's technical work is performed by the National Measurement Laboratory, the National

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'Headquarters and Laboratories at Gaiihersburg, MD. unless otherwise noted; mailing address

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'| ocaicd at Boulder. CO, with some elements at Ciaithersburg, MD.
Technical Digest—Symposium on Optical

Fiber Measurements, 1986

RESEARCH

INFORMATION CENTER

Digest of a Symposium sponsored by the

National Bureau of Standards

in cooperation with the

IEEE Optical Communications Committee

and the Optical Society of America

September 9-10, 1986

National Bureau of Standards

Boulder, Colorado 80303

Edited by

G.W. Day

D.L Franzen

Electromagnetic Technology Division

Center for Electronics and Electrical Engineering

National Engineering Laboratory

National Bureau of Standards

Boulder, Colorado 80303

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

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

Issued September 1986
5

Library of Congress Catalog Card Number: 86-600563

National Bureau of Standards Special Publication 720

Natl. Bur. Stand. (U.S.), Spec. Publ. 720, 158 pages (Sept. 1986)

CODEN: XNBSAV

U.S. GOVERNMENT PRINTING OFFICE

Washington: 1986

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

One might expect that, as a technology matures, associated measurement

problems would be solved and research in those areas would diminish. That

view is supported by examination of the programs for the annual Optical Fiber

Communications Conference (OFC) and European Conference on Optical Communications (ECOC) which have shown a steady decline in the percentage of papers on measurement topics over the last 10 years. It is not supported by our experience with the Symposium on Optical Fiber Measurements. Rather, it

appears that the need for a specialized meeting on the topic of measurements

continues to grow.

We believe that the strength of this Symposium is best indicated by the

number and quality of contributed papers. The number of papers submitted to

the committee has grown by roughly 20 percent with each succeeding Symposium.

Unfortunately, the time available for presentation could be increased at only

half that rate. This year, only about 50 percent of the papers received by

the program chairman could be accepted. Those that were accepted are diverse

in both subject and origin. In only two technical areas, chromatic dispersion

and joint characterization (including mode field diameter measurements), was

it possible to schedule a full session on a single topic. In origin, there

are 21 organizations in 9 countries represented in the program. Over 40 percent of the papers are from outside the U.S.

This Digest contains the text of all 34 papers presented (5 invited and

29 contributed). In addition to these formal presentations, two informal

workshops were held. One, on "Components and Measurements for LAN's and Wideband Premise Networks" was moderated by Allen Cherin of AT&T Bell Laboratories. The other, "Measuring the Future—Strategies for Upgradabi 1 i ty" was moderated by Paul Reitz of Corning Glass Works. There is no written record of

the workshops.

We continue to be indebted to the IEEE Optical Communications Committee

and the Optical Society of America for their assistance in organizing this

Symposium. The members of this year's committee, listed on the following

page, have been especially helpful in providing timely reviews of manuscripts

and other suggestions on the program.

G. W. Day

D. L. Franzen

Boulder, Colorado

September 1986

i i i
SYMPOSIUM COMMITTEE

D. L. Franzen, NBS, General Chairman

G. W. Day, NBS, Program Chairman

A. H. Cherin, AT&T Bell Laboratories

R. L. Gallawa, NBS

P. Kaiser, Bell Communications Research

F. P. Kapron, ITT

T. Ito, NTT

D. B. Payne, BTRL

P. R. Reitz, Corning

0. Szentesi, Siecor

i v
CONTENTS

Page

Preface iii

Conference Committee iv

Status of single-mode fiber measurements (invited)

William T. Anderson; Bell Communications Research 1 Study of algorithms used to fit group delay data for single-mode

optical fibers

W. A. Reed, D. L. Philen; AT&T Bell Laboratories 7 Chromatic dispersion in optical fibers: Comparison of

measurements with models

Leslie Button, Walter F. Love, Robert M. Hawk; Corning

Glass Works 11

Accuracy of phase shift technique with LEDS for measuring

total dispersion in singlemode fibers

L. Bosselaar, 6. Kuijt, J. F. van Luijk, P. Matthijsse

Philips Glass 15

High speed and high accurate single-mode fiber dispersion

measuring equipment using modified interferometric method

Y. Namihira, Y. Iwamoto; KDD 19

Chromatic dispersion measurements of single mode fibres

in the 1.3 ym and 1.55 ym wavelength range by the

wavelength temperature tuning technique

Finn Mogensen, Hans Damsgaard, Ole Hansen; NKT Elektronik 23

Measurement of dopants in optical fibers by Raman scattering

Dan L. Philen; AT&T Bell Laboratories 27

Utilizing cladding mode propagation to determine the refractive

index profile of an optical fiber

W. H. Hatton, E. L. Buckland, M. Nishimura; Sumitomo

Electric Research Triangle 31

Determination of refractive-index distribution from near-field

measurements in polarization maintaining fibres

G. Coppa, P. Di Vita, M. Potenza; CSELT 35

Prism coupling technique: A novel method for measurement of

propagation constant and beat length in single mode fibers

K. Thyagarajan, M. R. Shenoy, M. R. Ramadas; Indian

Institute of Technology 39

Characterization of optical fiber splices in the field (invited)

Reinhard Engel, Siemans; Everett McNair, Siecor 43

Consistency of mode field diameter definitions for fibers

with Gaussian and non-Gaussian field profiles in

different measurement domains

W. T. Anderson, J. P. Kilmer; Bell Communications Research 49

A new technique for mode field diameter determination by

transverse offset measurements

M. Calzavara, G. Coppa, P. Di Vita; CSELT 53

v
Joint losses as a function of offset and mode field

diameters for dispersion-shifted and standard fibers

and fiber combinations

V. Shah, L. Curtis, W. C. Young; Bell Communications Research 57

A novel single-mode fiber splice alignment and loss measurement

using local detection

Ian A. White, Paul D. Ludington; AT&T Bell Laboratories 61

Fusion splicing of dissimilar fibers —A comparison of mode

field diameter and cross correlation loss predictions with

experimental results

W. T. Anderson, A. J. Johnson, J. P. Kilmer, E. A. Thomas;

Bell Communications Research 65

Stress-induced increases in interference effects at fiber joints

and a method to eliminate interference effects in optical

fiber measurements

W. C. Young, V. Shah, L. Curtis; Bell Communications Research 69

Practical one-way-loss range for OTURs

M. C. Brain, E. A. Cottrell; British Telecom Research

Laboratories 73

A new technique for the relative measurement of scatter

levels in single mode fibers

M. E. Fermann, S. B. Poole, D. N. Payne, F. Martinez;

Southampton University 77

Mode field diameter effects on OTDR splice measurements

Felix Kapron, Carrie Kozikowski, Ricky Crotts;

ITT Electro-Optical Products Division 81

Multimode fiber systems characterization

I. P. Vayshenker, D. R. Hjelme, A. R. Mickelson;

University of Colorado 85

Single-ended bandwidth measurement for multimode optical

fiber using far end pulse reflection

Shin-ichi Furukawa, Katsuya Yamashita, Mitsuhiro Tateda,

Yahei Koyamada; NTT 89

Effective fiber bandwidth in LED based systems

Dieter Schickentanz, William S. Jackman; Siecor 93

Measuring 825 nm transmission on 1300 nm single-mode fibers

Catherine M. Ragdale, Felix P. Kapron; ITT 97

Characterization of sources and detectors for optical

fiber systems (invited)

T. Ito, K. Kurumada; NTT 101

Characterization of specialty fibres and components (invited)

D. N. Payne, S. B. Poole; Southampton University

M. P. Varnham; British Aerospace

R. D. Birch; Southampton University 107

Measurement requirements for waveguide electro-optic

devices (invited)

R. C. Alferness; AT&T Bell Laboratories 115
Practical upper limits to cutoff wavelength for different single

mode fiber designs

L. Wei, C. Saravanos, R. S. Lowe; Northern Telecom

Canada Limited 121

New measurement technique for measurement of microbend

losses in single mode fibers

A. 0. Garg, C. K. Eol 1 ; Siecor 125

Predicting microbending losses in single-mode fibers

Neil Kamikawa; Naval Ocean Systems Center

Ching-Ten Chang; San Diego State University 129

On the numerical aperture of a single-mode fiber

D. N. Christodoulides, L. A. Reith, M. A. Saifi;

Bell Communications Research 133

Optical fibre far field ESI determination using the

equalisation wavelength

A. C. Boucouvalas, S. C. Robertson, J. Walker;

GEC Research 137

Impact of reflection on the performance of single-mode

fiber optic systems

Shuenn Jyi Wang; Bell Communications Research 141

Spurious edge elimination for accurate glass geometry

measurements

Devon R. Mcintosh, David A. Hall; Corning Glass Works 145

Index of Authors 149

Except where attributed to NBS authors, the content of individual sections

of this volume has not been reviewed or edited by the National Bureau of Standards. NBS therefore accepts no responsibility for comments or recommendations therein. The mention of trade names in this volume is in no sense an endorsement or recommendation of the National Bureau of Standards.

vii

STATUS OF SINGLE-MODE FIBER MEASUREMENTS

by William T. Anderson

Bell Communications Research

1. INTRODUCTION

Single-mode fiber manufacture has advanced dramatically over the past several years. Today, typical

fibers approach the performance levels of the champion results reported only five years ago, with the

attenuation approaching the theoretical minimum levels dictated by Rayleigh scattering. Consistent

with the advance of the manufacturing state of the art, single-mode fiber measurements have advanced

in ease and accuracy, and state of the art measurements which could be made only in the laboratory in the past are now being made in the factory and in the field. However, with these advances in the

measurement state of the art, new problems have been discovered. In this paper, the current state of the art will be briefly reviewed with emphasis on measurements or specification issues where controversy

still exists. In the following sections, the four major measurements will be covered: attenuation,

dispersion, mode field diameter, and cutoff wavelength.

2. ATTENUATION

The measurement of attenuation is the oldest and, conceptually, the simplest of the single-mode fiber measurements. Three measurement techniques are commonly used. The most familiar of these is the

cutback technique. Typically, a tungsten-halogen lamp and monochromator provide a low-power but

wavelength-tunable source, and a lock-in amplifier is used for synchronous detection of the output of an

unbiased PIN photodiode. The repeatability of the cutback technique is dependent primarily upon the

ability to repeatably couple the fiber output to the detector. Standard deviations of .01 to .04dB and

the dynamic ranges of 20 to 35dB are typical for this technique.

For installed spans or connectorized cables, a destructive technique such as cutback is inappropriate. In these cases, the attenuation by substitution or insertion loss techniques are commonly employed. The

launch and detector coupling are made through connectors, and unlike the cutback technique, both

launch and detect coupling must be accurately reproducible. Also, for an installed span the source and

detector are separated so that either two detectors or a source with excellent long term stability are

required. A systematic bias can be introduced by using two detectors which are not accurately cross- calibrated or by using fiber jumpers whose mode field diameters differ significantly from those of the

fibers in the installed span. These factors clearly make the characterization of installed spans less accurate than the characterization of fiber and cable using the cutback technique.

A third commonly used attenuation measurement technique is backscatter, which employs an optical

time domain reflectometer (OTDR). Unlike other measurement techniques, OTDRs can locate faults, measure splice losses, and resolve the attenuation coefficient throughout the installed span. OTDRs are

also single-ended, so the measurement of installed spans requires only one craftsman rather than the two

required by the substitution method. However, OTDRs have two significant disadvantages. First, the

resolution of even the best and newest units is on the order of 100m, making it impossible to resolve the

first one or two splices in an installed span. Second, OTDRs do not measure loss directly, but infer loss from the measured backscatter of the fiber. Differences in the backscatter capture coefficient along the span can produce inaccurate loss estimates, especially for mismatched fibers, and reports of splice gain are common. Averaging two bidirectional measurements can eliminate this effect, but only at the cost

1
of doubling the time and effort required to make the measurement.

While the measurement of attenuation is relatively straightforward, the specification of attenuation is complicated by the uncertainty of the wavelength of the sources. Based upon the typical wavelength

ranges of currently used laser diodes, a wavelength "window" (e.g., 1285 to 1330nm) is often used as the

basis for a specification on the allowable variation in the attenuation with wavelength (typically

O.ldB/km). However, with the advent of LEDs for subscriber loop applications and the possibility of wavelength division multiplexing in the future, a rigid window specification appears to be inadequate.

Since the level of OH contamination in the fiber is typically the major limitation to the width of the

low-loss window, a specification on the maximum loss at 1390nm may be more efficient than a

multiplicity of windows for the multiplicity of possible sources. The specification and measurement of

this "water peak" needs further study.

3. DISPERSION

The great information transmission capacity of single-mode fibers which makes them so attractive as a medium also makes the measurement of the information capacity very difficult. Since the major

limitation in capacity for systems operating below tens of Gbit/sec is chromatic dispersion, indirect

measurement techniques for determining the information transmission capacity by characterizing this change in group velocity with wavelength are often used. Two similar techniques are frequently

employed: measuring the transit time of pulses at different wavelengths and measuring the phase shift of a modulated carrier at different wavelengths. The former approach may use a fiber raman laser' 1

! or multiple laser diodes'2

' as the optical source, while the latter approach may employ either multiple laser diodes'3

' or an LED and monochromator'4

'. The directly measured delay data is usually too noisy or too

discrete to allow numerical differentiation directly with acceptable accuracy, so the delay data is usually fitted to an analytical function, which can then be differentiated directly. However, there is some doubt about the validity of the three term sellmeier expansion which is most often used as a

functional form for fitting delay data'5

'. This is especially true of dispersion-shifted and -flattened fibers,

for which no functional form is known a priori.

If instead of measuring group delay, one measured the change in group delay between two closely spaced

wavelengths, then this differential delay would be an accurate approximation to the dispersion if the wavelength spacing were sufficiently small. A similar concept is used in differential spectroscopy. Two

means of measuring differential delay are shown in figures 1 and 2. In the first method, an LED is used

as a broad spectral source, and two closely spaced wavelengths are selected and detected. The phase

difference is directly measured, and an accurate approximation to the dispersion is obtained without any

assumptions about its spectral shape (except, of course, that the dispersion is slowly varying with

wavelength). A second method employs a low frequency wavelength modulation on the output of the high frequency amplitude modulated LED, for instance by wobbling a monochromator grating'8

'. This

introduces a phase modulation in the output of the detector due to the variation in transit time in the

fiber with wavelength, and this phase modulation is detected synchronously from the quadrature output

of the high frequency mixer. This method offers more flexibility in chosing the spectral width and

wavelength modulation width than does the other differential modulation technique.

The direct measurement of differential delay promises to eliminate the uncertainties of data fitting,

although, in principle, similar benefits can be derived from the conventional delay measurement

approach if the number of wavelengths is sufficiently large and the measurement error is suitably small.
Increasing the number of measurement wavelengths is feasible with either the fiber raman laser or the

LED/monochromator, but not with multiple laser diodes. It may be possible to reduce the size and

complexity of fiber raman sources by using non-silica raman fibers, which can be pumped with

semiconductor lasers'7

!

The specification of dispersion has been a subject of recent controversy. The usual approach of

specifying the maximum dispersion in a fixed wavelength window, similar to the specification of the

wavelength variation of attenuation, is too inflexible, to permit efficient system design using a

multiplicity of possible sources, such as LEDs. For conventional step-index fibers, the dispersion is approximated by a sellmeier expansion. While this approximation may be too inaccurate for the

measurement of dispersion, it may be sufficiently accurate for the specification of dispersion. The zero

dispersion wavelength and the slope of the dispersion at the zero dispersion wavelength, used with the

sellmeier approximation, have been proposed as the basis for the specification of dispersion. This

method is more flexible than a window specification and permits the calculation of the dispersion over any desired wavelength window. However, there are questions about the accuracy and consistency of dispersion measurements made using the different techniques, and a round-robin series of measurements

are planned to resolve this issue. For dispersion-shifted and -flattened fibers, for which no general

functional form is known, a different method of specification is needed.

4. MODE FIELD DIAMETER

In contrast to the difficult measurement of dispersion, the measurement of the mode field diameter is quite easy to implement. Perhaps it is too easy since there are so many proposed measurement

techniques, and the challenge facing the industry is to ensure that all measurement techniques give

consistent results for all fiber designs. Five techniques are currently being considered by the EIA and

CCITT for standardization:

1. Scanning the fiber near fields by focusing a greatly magnified image of the fiber core either onto an

infrared vidicon or onto the plane of a scanning small-area detector'8

'.

2. Measuring the transmission through a splice as a function of offset, from which the two

dimensional autocorrelation of the near fields can be obtained'9

'.

3. Measuring the fiber far fields any of three ways:

a. Scanning a small area detector in the far field'

10

'.

b. Measuring transmission through apertures in the far field as a function of aperture radius' 11

'.

c. Measuring transmission past a knife edge in the far field as a function of lateral position' 12

'.

All of these techniques have certain advantages, and there appears to be little chance that any one of them will be widely accepted as a standard, or reference, technique by which the others could be judged.

Therefore, consistency among the techniques is essential if the mode field diameter is to be used as a measure of the optical compatibility of two fibers.

The degree of consistency achievable with these five techniques is a direct result of the choice of a

definition for the mode field diameter. Originally, the gaussian approximation, proposed by Marcuse' 13'

to explain single-mode fiber splice loss, was applied to data generated by the various measurements, but,

since the fields of even a perfect step-index fiber are only approximately gaussian, small but systematic

differences among the measurement techniques were observed for quasi-step index fibers, especially at longer wavelengths' 14 "

15

'. Differences of greater than 10% have been observed for non-step-index

dispersion-shifted fibers'

16

'. Clearly, the gaussian approximation is not universally applicable.

3
An alternative definition based upon the second moment of the far fields, which is frequently called the "Petermann 2" definition! 17 " 18

', caIli j n principle, solve this problem. Since the five measurements are

related by several transforms to a good approximation, the second moment definition can be mapped

from one measurement domain to another, and, in theory, consistency can be maintained. For example,

the near fields, far fields, and autocorrelation are related by Hankel transforms. The far fields, knife edge scan, and variable aperture scan are related by Abel transforms. The knife edge scan and the

autocorrelation function are related by the Fourier transform. Consistency between three of the five techniques has been demonstrated, and a round-robin measurement of non-step-index fibers, which is needed to confirm the applicability of this definition to all measurement techniques, is planned.

A consistent definition and improved measurement accuracy could lead to a more accurate specification

of the mode field diameter. Since the specified mode field diameter is used to estimate the loss to be

expected when dissimilar fibers are spliced, a more accurate specification would give a more accurate

picture of the compatibility between fibers of different designs. Also, as splicing equipment and

techniques improve, a larger fraction of the splice loss is attributable to fiber parameter mismatch, so a

tighter tolerance on the variation of the mode field diameter may be appropriate.

5. CUTOFF WAVELENGTH

The concept of cutoff wavelength is deceptively simple. In theory, at wavelengths longer than the cutoff wavelength, the fiber supports only one mode (actually, a degenerate mode group). However, it is well known that the higher order modes become highly attenuated at wavelengths substantially below the

theoretical cutoff wavelength. By convention, the "effective" cutoff wavelength is defined as the

wavelength at which the higher order mode attenuation exceeds some chosen level and, therefore, higher

order modes cease to propagate significant distances. This is a universally accepted definition for telecommunication systems, where the distances between mode coupling perturbations, such as splices, are typically on the order of hundreds of meters, and the adverse effects of multimode propagation,

modal noise and distortion, can be avoided if the higher order modes are sufficiently attenuated. For

short distance applications, such as fiber sensors, there may be different requirements.

The attenuation level which defines the cutoff wavelength depends upon the measurement technique.

CCITT and EIA specify a measurement sample two meters long with a 28cm diameter loop' 19

'. The

cutoff wavelength can be measured using two possible methods, as shown in figure 3. First, cutoff is identified as the wavelength above which a small diameter (typically 3 to 5 en) loop changes the attenuation by less than O.ldB (the bend reference method). Alternatively, cutoff is identified as the

wavelength above which the transmission through the single-mode sample, when normalized by the transmission through a multimode reference fiber, approaches a linear region to within O.ldB (the

multimode reference technique). A round-robin measurement has verified that these two methods give

approximately equal values for the cutoff wavelength'20

'. Either of these methods equate cutoff with

approximately 20dB of attenuation for the higher order mode, although this attenuation may be

attributable to either the sample length or the bend, depending upon the fiber design.

These measurement conditions give a convenient and repeatable measurement, but the actual

deployment conditions of the fibers will be different from those in this measurement. Since both fiber length and bending significantly alter the measured cutoff wavelength'21 " 22

', relating measurements made

using this arbitrarily chosen length and bend condition do not relate simply to the design of telecommunications systems. To avoid modal noise and distortion, the attenuation of the higher order

4
mode must be large over the shortest distance between mode coupling perturbations, such as splices.

While this distance is typically kilometers, it can be substantially shorter for jumpers and pigtails, for stub cables, or under certain difficult installation situations or for repair sections. A measurement

method for determining the cutoff wavelength of cabled fiber, for which the user would specify a sample

length and deployment condition consistent with his application, is under consideration. While the

cable specification can be less conservative than the fiber specification, there is not universal agreement

that a less conservative specification is desirable, especially when using LEDs with single-mode fibers. A

better understanding of future applications in the subscriber loop is needed.

8. CONCLUSIONS

Single- mode fiber measurements continue to be developed and improved at a rapid pace. However, the

performance of fibers and splicing equipment has also improved, and greater accuracy is required to characterize fibers and splices. Because of the high cost of installing, splicing, and testing single-mode

fiber cables, their potentially high capacity is a compelling advantage since upgrading the capacity of an existing span may avoid additional construction. However, exploitation of this high capacity in the

future or initial operation with LEDs may require utilizing fibers in wavelength regions which are not

presently used. The uncertainty of possible future uses of the fiber complicates the measurement and

specification of attenuation, dispersion, and cutoff wavelength. This uncertainty can only be resolved

when the technology of optical systems, and especially that of the optical sources, is mature.

REFERENCES

1. L. G. Cohen and C. Lin, Applied Optics, Vol. 16, p. 3136, 1977.

2. C. Lin et al, Bell Syst. Tech. J, Vol. 62, p. 457, Feb. 1983.

3. K. Tatekura et al, Tech. Digest, Symposium on Optical Fiber Measurements, 1984, P- 119.

4. B. Costa et al, Electron. Lett., Vol. 19, p. 1074, Dec. 1983.

5. F. P. Kapron and T. C. Olson, Tech. Digest, Symp. on Optical Fiber Measurements, 1984, P- HI6. A. G. Barlow and R. S. Jones, to be published.

7. H. Takahashi et al, Optics Letters, Vol. 11, No. 6, p. 383, June 1986.

8. Y. Murakami et al, Applied Optics, Vol. 18, p. 1101, 1979.

9. J. Streckert, Opt. Lett, Vol. 5, p. 505, Dec. 1980.

10. K. Hotate and T. Okoshi, Applied Optics, Vol. 18, p. 3265, 1979.

11. E. Nicolaisen and P. Danielsen, Electron. Lett., Vol. 19, p. 27, 1983.

12. W. G. Otten and F. P. Kapron, Tech. Digest, OFC 86, Paper TUL30.

13. D. Marcuse, Bell Syst. Tech. J., Vol. 56, P. 703, 1977.

15. W. T. Anderson, IEEE JLT, Vol. LT-2, p. 191, April 1984.

15. D. L. Frantzen and R. Srivastava, IEEE JLT, Vol. LT-3, No. 5, Oct. 1985, p. 1073.

16. W. T. Anderson et al, measurement comparison to be published.

18. K. Petermann, Electron. Lett., Vol. 19, p. 712, 1983.

18. C. Pask, Electron. Lett., Vol. 20, p. 144, 1984.

19. CCITT Recommendation G.652 and EIA RS-455-80.

20. D. L. Frantzen, IEEE JLT, Vol. LT-3, No. 1, p. 128, Feb. 1985.

22. W. T. Anderson and T. A. Lenahan, IEEE JLT, Vol. LT-2, p. 238, June 1984.

22. V. S. Shah, Tech. Digest, Symposium on Optical Fiber Measurements, 1984, P- 7.

5
LEO

TEST

SAMPLE]

SPECTROMETER

0 HF

NETWORK <J> , ANALYZER <j> 2 ATTENUATOR

Figure 1- Differential Delay Dispersion Measurement

MONOCHROMATOR t

DETECTOR

LP

PHASE METER

ref

out

In

LF LOCK-IN

out

Figure 2- Double Modulation Dispersion Measurement

OVERFILLED

LAUNCH

(sample) y-v • POWER P. (\)

POWER P

2a (\)

POWER P

2b (\)

10 LOO

Figure 3- Cutoff Wavelength Measurement

6
Study of Algorithms Used to Fit Group Delay Data for Single-Mode Optical Fibers

W. A. Reed

AT&T Bell Laboratories

Murray Hill, New Jersey

D. L. Philen

AT&T Bell Laboratories

Norcross, Georgia

Introduction

The wavelength(s) of zero dispersion (X 0 ), and the slope(s) of the dispersion (S g ) at X,, are

critical parameters in specifying the transmission properties of single-mode lightguides. The

procedure used to calculate X„ and S0 is to measure the relative group delay in the wavelength

region of interest, fit an analytic function to the delay data, differentiate the function and

calculate the dispersion using the fitted coefficients, calculate X„ and S0 either from the

coefficients directly or by interpolation.

Although several analytic functions have been used to fit the group delay data 1

, the EIA is proposing a three term Sellmeier equation for normal unshifted (X„ « 1.3 /xm) fiber primarily

because \ 0 and S0 may be calculated directly from the fitted coefficients. As single mode fibers are designed for operation at 1.55 fim (dispersion shifted), operation at

1.55 fim with reduced S0 (dispersion flattened), and with multiple dispersion zeros, the question

arises whether the three term Sellmeier equation is adequate in all cases and, if not, is there a

single equation which will satisfactorily fit the group delay of all fiber designs.

In this paper we present the results of fitting five different algorithms to group delay data for four different fiber designs. The group delay data used was both measured experimentally and

calculated using computer aided modeling.

Description of the Algorithms

The algorithms studied were selected because they are currently in use and have been studied

by Kapron1 for standard unshifted fiber. The algorithms are as follows:

Three Term Sellmeier (S3)

r{\)= CiX"2+ c 2 + c 3X2 (1) D{\) = -2c!X-3 + 2c 3X (la)

The data can be fitted using a linear least squares (llsq) fit; X„ = (c l /c3)

1 /i and S0 = 8c 3

.

Five Term Sellmeier (S5)

r(X) = CjX-4 + c 2X"2 + c 3 + c 4X2 + c 5X" (2) D{\) = -4c

^- 2c 2X~3 + 2c 4X + 4c 5X

3 (2a)

The data can be fitted using a llsq fit; X„ and S0 are determined from interpolating the

calculated dispersion.

7
Three Term Power Series (P4)

t(X) = c 2 + c 3(X - e ,)2 + c 4 (X - Cl ) 3 (3) D{\) = 2c 3(X - Cl) + 3c 4(X - Cl ) 2 (3a)

Although the data can be fitted using a llsq fit, to preserve the significance of Cj and c 3 , a nonlinear least squares (nllsq) procedure was used. In this form \ 0 = c l and 5,, = 2c 3

.

Four Term Power Series (P5)

t(X) = c 2 + c 3(X - Cl ) + c

4 (X - Cl ) 2 + c 5(X - Cl ) 3 (4) D(X) = c 3 + 2c 4(X - Cl ) + 3c s(X - c,)2 (4a)

We fitted the delay data to the above equation using a nllsq fit although a llsq fit would also work using four coefficients and removing c l

. In either case X^ and S0 must be obtained by

interpolating the calculated dispersion.

Log Term (Ln)

r(X) = c 2 - c 3Cl2(^l - (5)

P(X) = c 3(-^)2 (X- Cl ) (5a)

The data is fitted using a nllsq fit; X„ = c : and S0 = c3

. For each equation that required finding X„ and S0 by interpolation, X,, was found approximately

and the dispersion calculated for AX = 0.0001 //m around the approximate zero(s). Linear

interpolation was then used between adjacent points to calculate X,, and S0

. With the

availability of inexpensive, fast microprocessors, any advantages of a linear vs. a non-linear fit or calculating X„ and S0 from the fitted coefficients is negligible.

Fiber Types Studied

Fibers exhibiting four basic types of dispersion were studied: a) standard unshifted

(X„ !=» 1.3 Ltm) fibers with both matched and depressed claddings; b) dispersion shifted

(X,, ~ 1.55 fim) fibers with triangular cores and both matched and depresses claddings; c) dispersion flattened (X 0 ~ 1.55 Lim, S0 reduced) fibers with triangular cores and depressed and

multiple claddings; d) dual zero fibers of the quadrupedal clad design2

. For each fiber type, the

group delay was measured experimentally, and calculated using a computer aided modeling

program. This was done to test the fitting procedures on data with and without "noise".

Results

Tables I-IV summarize the results of fitting the five equations to the modeled data for the four

fiber types. Fit indicates the algorithm used, npts is the number of points used in the fit, std. dev. is the mean square deviation, \ 0 is the zero disperion wavelength in microns, and slope is the dispersion slope at X„ in ps/km-nm2

. The data used to generate Tables I-IV were also reduced to 6 points about X„ (AX = 0.01 //m) and re-fitted since several commercially available

dispersion test sets use six semiconductor lasers.

For standard unshifted fibers (Table I) all algorithms fit equally well in terms of X„ and S0 except Ln which is slightly poorer. Over the restricted range of 6 points all algorithms fit well although S5 gives the smallest std. dev. When the experimental data is used the same

conclusions are reached.

8
For the dispersion shifted designs (Table II) all five algorithms find X 0 correctly but S3 and Ln

have slightly lower S0

, s and higher std. dev. When the number of points is reduced to 6 all algorithms work equally well. When the experimental data is used, S3, P4 and Ln give slightly

less reliable values for X,, and Sg with S5 and P5 giving consistently better results. The results for the dispersion flattened design (Table III) are similar to those for the dispersion

shifted design except that the error in S0 becomes magnified for S3 and Ln. The group delays for the dual dispersion-zero design presents a problem for all of the

algorithms. As can be seen in Table IV, S3 and Ln are unable to find two X^'s. Although P4,

P5 and S5 all find two X„'s, none finds both values accurately. Of the three however, S5 comes

the closest to the correct results.

Summary

Based on the five algorithms tested in this study, we find that the three term Sellmeier (S3) is very adequate for fitting group delay data for standard unshifted fibers (X,, = 1.3 fim).

However, in terms of a more general fitting equation that is applicable to a wide class of fiber types, we find that the 5 term Sellmeier equation (S5) gives consistently better results for all of the fiber designs tested using both experimental and modeled data. The 3 term Sellmeier (S3)

and log (Ln) algorithms are inferior as a general equations. With the availability of inexpensive, fast microprocessors, the questions of linear and non-linear fits, and the ease of finding \ 0 and S0 from the fitted coefficients is not a valid concern.

References

[l] Felix P. Kapron and Tom C. Olson, "Accurate Specification of Single-Mode Dispersion

Measurements", NBS Symposium on Optical Fiber Measurements, 1984, pg. 111.

[2] L. G. Cohen, W. L. Mammel, and S. J. Jang, "Low-Loss Quadruple-Clad Single-Mode

Lightguides with Dispersion Below 2 ps/km-nm over the 1.28 /zm-1.65 ^m Wavelength

Range," Elect. Lett. 18, 1023-1024 (1982).

9
Table I Model Standard Unshifted Fiber (X, = 1.3113, S0 = 0.0851)

fit npts std. dev. slope

S3 21 0.0279 1.3113 0.0851

S5 21 0.0000063 1.3113 0.0851

P5 21 0.111 1.3112 0.0857

P4 21 0.111 1.3112 0.0857

Ln 21 4.20 1.3105 0.0844

Table II Model Dispersion Shifted Fiber (X, = 1.5531, S0 = 0.0589)

fit npts std. dev. K slope

S3 16 0.749 1.5529 0.0571

S5 16 0.000882 1.5531 0.0589

P5 16 0.00262 1.5531 0.0589

P4 16 0.00233 1.5531 0.0589

Ln 16 1.77 1.5528 0.0560

Table III Model Dispersion Flattened Fiber (X„ = 1.5622, S0 = 0.0307)

fit npts std. dev. slope

S3 16 6.99 1.5621 0.0386

S5 16 0.00514 1.5622 0.0306

P5 16 0.0200 1.5625 0.0309

P4 16 0.0198 1.5625 0.0308

Ln 16 5.45 1.5621 0.0377

Table IV

Model Dual Zero Fiber (X, = 1.4597, 1.6063, S0 = 0.0336, -0.0562)

fit npts std. dev. K slope

S3 31 2360 1.4927 0.00931

S5 31 12.4 1.4606 0.0380

S5 31 12.4 1.6024 -0.0516 P5 31 91.1 1.4520 0.0386

P5 31 91.1 1.5996 -0.0386

P4 31 91.1 1.4520 0.0386

P4 31 91.1 1.5997 -0.0387 Ln 31 2310 1.4914 0.00965

10
CHROMATIC DISPERSION IN OPTICAL FIBERS: COMPARISON OF MEASUREMENTS WITH MODELS

Leslie Button, Walter F. Love, Robert M. Hawk, Research and Development Division,

Corning Glass Works, Corning, New York 14831

INTRODUCTION

Chromatic dispersion is an important broadening mechanism for time dispersion

in optical fibers. It is the limiting factor in LED-multimode systems and in

single-mode systems apart from modal noise considerations. In this paper we present a comparison of time dispersion transmission measurements and refractive

index data and models for both graded-index multimode fibers and single-mode

fibers

EXPERIMENTAL METHOD

The time delay, t, was measured for fiber lengths, L, of 2 to 5 km over the

1 2 wavelength range, 1.1 to 1.7 urn using a fiber Raman laser system. ' These data 2-2 were fitted to t=A+BX +CX for multimode fibers and unshifted single-mode fibers,

TM

SMF ; whereas, T=an+a q X+a

9 XlnX was used for the dispersion-shifted single-mode

TM fibers, SMF/DS , where X is the wavelength of light. The chromatic dispersion,

1 dx

D, is readily obtained as D = — -jy. A mode-scrambled launch is used to fully

excite all modes in the multimode fibers. The precision in all cases is typically

~1 nm.

3 The RNF index profile measurement uses a He-Ne laser operating at 633 nm. The finite laser spot size gives an instrument resolution of 0.6 urn. The measured

index profile serves as input data to a computer program which numerically solves

the scalar wave equation using the finite element method. Refractive index is

required as a function of wavelength, viz., n(X), and we have used the published

work of

fibers.

4 5 6 work of Kobayashi , Fleming , and Shibata in our comparisons for single-mode

MULTIMODE FIBER RESULTS

Measurements of the zero dispersion wavelength, X , are plotted versus an effective A value based on far-field NA measurements for high-bandwidth graded

index fibers and are shown in Fig. 1. The zero-doped value was obtained from

11
Malitson's index measurements of fused silica.'

7 There is a high degree of

correlation and a linear dependence up to approximately 2%. Thus, X

q is shifting

toward longer wavelengths with increasing germania doping.

The chromatic dispersion, D, has been calculated using the expression for x

g obtained by WKB analysis for a-model index profiles. Assuming equal mode

9 excitation and a parabolic index profile, we obtain in agreement with Adams ,

2 M

1 + M9 n1 Ap

D= 3 +3^- (1)

X d

2 n. where the material dispersion, M. = x— and the subscript 1 and 2 refer to

1 c dXZ core and cladding. The profile dispersion parameter, p, is given by

p - ^ and c equals the speed of light.

The profile dispersion term can be evaluated near 1300 nm, where p - 0.05,

-1 -1

for A = 0.01 to give n..Ap/3cX = 0.63 psec nm km . At a slope of 0.09 psec

-2. -1 . . -, . \L nm km (Ge02~Si02), this represents a 7 nm correction. Since the material

dispersion varies linearly with wavelength near Xq, Eq. 1 can be reduced to:

\> " V

2>

+ J

< X

o (1)

- X

o

<2)

> - 7 nm < 2 >

11 12 • „ Using the measurement results of Payne and Hartog ' for X

q versus mole X

Ge02-doped silica, we compute for 12 mZ, corresponding to 0.20 transmitted NA,

X

q

= 1274 + 2/3 (1340-1274) - 7 = 1311 nm. This agrees very well with our result

shown in Fig. 1.

SINGLE-MODE FIBER RESULTS

Schematic diagrams of the index profiles for unshifted, SMF , and

TM dispersion-shifted, SMF/DS , single-mode fibers are given in Fig. 2. The actual

profiles for five of the former and four of the latter were measured at each end

by the RNF method. These were entered as input data to a finite element solution

of the scalar wave equation along with published index data.

The difference between the long-length measurements and the index-calculated

12
values for X

q

, viz., SX

q

, are given in Table 1. The standard deviation about the mean value of SX is denoted by a . The best agreement is closest to Fleming for

TM

0 m SMF . This is likely to shift towards the less dispersive index data of

Kobayashi with deconvolution of the input spot size. This will be discussed in

TM more detail later. For SMF/DS , the average agreement is better although there

is more imprecision, viz., a is larger. This is understandable since the

waveguide dispersion is dominant and small variations in index profile measurement

give significant variations in X .

A more detailed view of the difference between measured and calculated

TM -" ' dispersion is given in Fig. 3 for SMF fibers. The shaded regions encompass ± la

values.

SUMMARY

We have compared long-length transmission measurements of chromatic

dispersion for high-bandwidth graded-index multimode fibers and single-mode fibers

with calculations based upon published index and dispersion data. We find very

good agreement with University of Southampton results for the multimode case. For

unshifted single-mode fibers, the agreement with calculation is best for Fleming's

index data whereas for dispersion-shifted single-mode fibers agreement is good for

all published index data considered, but the precision is less in this case.

REFERENCES

1. L. G. Cohen and Chinlon Lin, IEEE J. Quantum Electronics QE14 , 855 (1978).

2. W. F. Love, Proceedings of Sixth European Conference on Optical

Communication, p. 113, York, U.K. (1980).

3. W. J. Stewart, Proceedings of Conference on Integrated Optics and Optical

Communication, C2.2, Japan, 1977.

4. S. Kobayashi, S. Shibata, N. Shibata, and T. Izawa, Proceedings of Conference

on Integrated Optics and Optical Communication, B8-3, Japan, 1977.

5. J. W. Fleming, Electronics Letters 14, 326 (1978).

6. N. Shibata, S. Shibata, and T. Edahiro, Electronics Letters 17, 310 (1981).

7. I. H. Malitson, J. Opt. Soc. Am. 55, 1205 (1965).

8. R. Olshansky and D. B. Keck, Applied Optics 15, 483 (1976).

9. M. J. Adams, D. N. Payne, F. M. E. Sladen, and A. H. Hartog, Electronics

Letters 14, 64 (1978).

10. F. M. E. Sladen, D. N. Payne, and M. J. Adams, Electronics Letters 15, 469

(1979).

11. D. N. Payne and A. H. Hartog, Electronics Letters 13, 627 (1977).

12. M. J. Adams, D. N. Payne, F. M. E. Sladen, and A. H. Hartog, Electronics

Letters 14, 703 (1978).

13
INDEX PROFILE

Fig. 1 Multimode fiber Aq versus Fig. 2 Ideal index profiles for

effective A from far-field. single-mode fibers.

3

X (/j.m)

Fig. 3 Comparison of transmission

measurements versus calculations for various

index data.

14
ACCURACY OF PHASE SHIFT TECHNIQUE WITH LEDS FOR MEASURING TOTAL

DISPERSION IN SINGLEMODE FIBERS

L. Bosselaar, G. Kuijt, J.F. van Luijk, P. Matthijsse

Telecommunication fiber development - Philips Glass

P.O. Box 218 - 5600 MD Eindhoven - The Netherlands

1. Introduction

The phase shift technique for measuring total dispersion coefficient has

been described in several publications 1

, 2 . Main advantages of using LEDs

are the compactness of the set-up and the possibility of measuring at

many wavelengths with two sources only. In this summary the main effects

determining the accuracy of this method are described. The results

presented have been obtained as follows : A theoretical

delay-time-versus-wavelength curve was calculated for depressed-cladding

and for dispersion- flattened singlemode fibers^. This curve was disturbed with random noise and systematic errors. The disturbed curves were fitted through a limited number of points at several wavelengths

with different types of curve-fits. The resulting dispersion

coefficients at specific wavelenghts were compared with the calculated

exact values

.

2 . Random phase noise

This noise may have different sources, and will lead to a spread in the

total dispersion measured. It was found that the spread in total

dispersion coefficient increases linearly with the amplitude of the

random noise for all the curve-fits. Tanaka2 has found similar results.

3. Temperature changes

The group velocity of light in the fiber is temperature dependent^. The

temperature coefficient is about 30 ps km~l K~l for primary coated

fibers. A change in temperature of the fiber during a measurement will

consequently affect the total dispersion coefficient measured. For this error to be kept below 0.1 ps km~^nm- l, the temperature change during a measurement should be less than about 1 K, dependent upon the temporal

behaviour of this change (worst case; less than 0.1 K)

.

15
4. Number of wavelengths at which phase shift is measured

The effect on the accuracy of the number of wavelengths measured has

also been investigated by the method described in the introduction. The

spread in the total dispersion coefficient of a depressed cladding

singlemode fibre measured at 1330 nm (D1330) versus the number of points

(N) at which the phase is measured is typically :

N : 10 20 30 40

Spread in D1330 : 0.2 0.06 0.04 0.02

The table shows a nearly quadratic reduction in spread of D1330 with

increasing N.

5 . Spectral width of monochromator

The spectral width of the monochromator will be chosen as high as possible to increase the received power and dynamic range. On the

contrary the spectral width should be limited, for prevention of

systematic error in the total dispersion coefficient measured. This

systematic error has been calculated for two spectral widths of the

monochromator, assuming there is one LED with a central wavelength at

1255 nm and one at 1540 nm. Fig. 1 shows the calculated error for the

depressed-cladding singlemode fibre; fig. 2 shows this error for the

dispersion- flattened singlemode fibre.

rig. 1 LED 1 : 1255 nm ERROR IN TOTAL DISPERSION CAUSED BY THE LED 2 : 1540 nm SPECTRAL WIDTH OF THE MONOCHROMATOR

ps/nm km DEPRESSED -CLADDING SINGLEMODE FIBRE

. 15 nm

. 1

,-6 nm

.05 ——2 nm

t-r-^

-—3 nm —1 1-

-.05

-. 1 •

1200/ 1300 1400 1500 1600 nm

-. 15

16
Figures 1 and 2 prove that the error in total dispersion is less than

0.1 ps km- lnm~l in the normally used transmission windows for a spectra

width of the monochromator of 6 nm FWHM. As this error is systematic, i can be compensated for if some knowledge of the expected dispersive

behaviour of the fibre is present.

Fig. 2 LED 1 : 1255 nm ERROR IN TOTAL DISPERSION CAUSED BY THE LED 2 i 1540 nm SPECTRAL WIDTH OF THE MONOCHROMATOR.

ps'nm kit DISPERSION-FLATTENED SINGLEMODE FIBER

.05

6 nm

nm

3 nm

3 nm

-.05

1200 / 1300 1400 1500 1600

—

h

nm

-. 1

•

6. Type of curve— fit

Among others the following fits with different numbers of coefficients

have been tried for the disturbed delay-time-versus-wavelength points :

S(x)=sj+S2X2 +S3x~2+s4x^+S5X-

^ .... (material dispersion fit,

based on Sellmeier's equation

T(x)=ti+t2(x-xo)+t3(x-xo)2 + t4(x-xo)3 ... (general Taylor series)

The 5-coef f icients Taylor fit (T5(x)) and the 5-coef f icient material

dispersion fit (S5(x)) were acceptable. Best results however, were obtained from the 3-coeff icient material dispersion fit for the

depressed-cladding singlemode fibre only and the 4-coeff icient material

dispersion fit for both types of fibres. The use of more coefficients

caused oscillations in the dispersion curve.

17
The criteria for usefullness of a fit-curve were :

1) average deviation and spread of total dispersion coefficient for

1285, 1330 and 1550 ran

2) average deviation and spread of the wavelength of zero total

dispersion (A 0 ) 3) slope and its spread of total dispersion coefficient in \ Q

(depressed-cladding singlemode fibre only)

7. Conclusions

- Measures for temperature stabilisation of the measured fibre must be

taken into account.

- An increase in number of wavelengths at which the phase shift is

measured, has a more favourable effect on the total accuracy of the

measurement than an increase in precision of the measurement of the

phase shift at one wavelength.

- Errors due to the non-zero spectral width of the monochromator

normally are rather small.

- The 4-coef f icient material dispersion fit yielded the best results

of all fits tested for both depressed-cladding and dispersionflattened singlemode fibres.

- The total accuracy of this method can be estimated as better than

0.5 ps km~^nm~l for standard production lengths of fiber in the 1300

and 1500 nm windows. Measuring results will be presented at the

conference

.

References

1. B. Costa, D. Mazzoni, M. Puleo, E. Vezzoni, "Phase shift technique

for the measurement of chromatic dispersion in optical fibers using

LED's", IEEE journal of quantum electronics, vol. QE-18,no. 10,

October 1982.

2. S. Tanaka , Y. Kitayama, "Measurement accuracy of chromatic

dispersion by the modulation phase technique.", Journal of

Lightwave Technology, vol. LT-2 , no. 6, December 1984 p. 1040.

3. P.K. Bachmann et al . , "PCVD DFSM-FIBRES: PERFORMANCE, LIMITATIONS,

DESIGN OPTIMIZATION", 11th ECOC, VENECIA, 1985, Conf. Proc .

, P. 197.

4. L.G. Cohen, J.W. Fleming, "Effect of temperature on transmission in

1 ightguides" , The Bell system technical journal, vol. 58, no. 4, p. 945-951, April 1979.

18
HIGH SPEED AND HIGH ACUCURATE SINGLE-MODE FIBER DISPERSION

MEASURING EQUIPMENT USING MODIFIED INTERFEROMETRIC METHOD

Y. NAMIHIRA, Y. IWAMOTO, T. MURAKAMI* and S. AN*

K D D Research & Development Laboratories,

2-1-23 Nakameguro, Meguro-ku, Tokyo, Japan

(*) SANTEC Corporation, Kamisue Komaki, Aichi, Japan

INTRODUCTION In the conventional inter ferometr ic chromatic

dispersion measurement [1],[2],[3], the high speed measurement

and compact arrangement could not be achieved. However, using

modified inter ferometr ic method [4] with AC-modulat ion detection

technique of the interference signal and inter ferometr ic filter

instead of monochrometer , the high speed, high accurate and

compact arrangement for single-mode fiber chromatic dispersion

measuring equipment can be developed.

In this paper, the modified interf erometr ic dispersion

measurement in short length single mode fiber is presented.

MEASURING SYSTEMS Fig.l shows a schematic diagram of the high

speed and high accurate single-mode fiber dispersion measuring

equipment (SANTEC SDA-6000) using modified interferometr ic method

[4], and the appearance of the equipment is shown in Fig. 2. The

size of the equipment is about 905 (W) x 795(D) x 1100(H) (mm).

The specifications of the measuring equipment are summarized in

Table 1. As an optical source and detector, a 100-W halogen lamp

and a InGaAs-APD were used. And, as a wavelength selector, the

interf erometr ic filter (FWHM - 15 nm) was used to realize compact

measuring equipment without using monochrometer. The dispersion

equipment has a function of automatic control high speed signal

processing sub-systems [4]. The input light is fed through the

interf erometr ic filter, and then launched into the lead singlemode fiber. After passing through the fiber, the light is devided

into the input beamsplitter and launched into both arms of the

Mach-Zehnder interferometer. A multi-reflection mirror was used

in the interferometer reference arm in order to equalize the

optical paths of the test fibers. The length of test fiber

samples is about 1.300 (m) ± 5 (mm). A variable optical delay

19
line is placed in the reference arm of the interferometer. The

principle of dispersion measurement is based on the conventional

interf erometr ic method [1],[2], however, in our modified interfero

-metric measurement, in order to search the interference point in

a short time, the optical path of the refernce wave is modulated

by actuating corner cuve with PZT at 60^100 Hz sinusoidal waves

instead of optical chopper, and the modulated interference

visivility fringe can be obtained [4]. A typical peak search

pattern of the interference visivility fringe is shown in Fig. 3.

RESULTS & DISCUSSIONS The dispersion D =dT/d$\. is obtained by

least square fitting equation to be measured delay time T and

differentiating it with respect to wavelength 7l . Also, the

dispersion slope is obtained as dD/dfl.. The delay time and

dispersion characteristics of the segmented-core dispersion

shifted fiber (DSF) [5] as a function of wavelength are shown in

Figs.4-(a) and (b), respectively. The delay time data (x) at 26

different wavelengths were measured by the modified interferometric dispersion measurement [4]. As a fitting equation [6], [7],

2 the following three-term Sellmeier equation = A A. + B +

C ft

-2

and the five term-Sellmeier equation ~T = A 7L + B Ji? + C + D A.— 2—4 and the Napierian logarithm expression

= A + B X + C /1-lnA were used. Where, A, B, C, D, and E are

coefficient of the delay timeT fitting expressions. In the delay

time curve-fitting expresions, the coefficients from A to E can be

determined by RMS fitting to the experimental data. From Fig. 4-

(a), it is found that the fitting curves of ( ) an<3 T3

( ) agree well with the experimental results ( x ) rather than

that of T ( )• The comparison of the three fitting expressions

are summarized in Table 2. From Table. 2, it is found that the

meansquare fitness error and dispersion-slope of < . is great

different from that of ~Z

2 and "^3* It: -"- s considered that the

lower-order Sellmeier equation of "X^ is strongly dependent of

the material dispersion, however, in the case of the DSF, the waveguide dispersion becomes greater than the conventional dispersion

non-shifted fibers. Accordingly, the fitting expression of ( does not agree with the measured values. Meanwhile, the results

of T2 anci T3 indicate almost the same values. At present, it

is not concluded that which types of the fitting equation of ~T

20
and <* ^ are tne best fitting expression for the dispersion

shifted fibers or not, and accordingly, the further study of the

gruop delay fitting equations will be needed. Meanwhile, in order

to confirm the measuring accuracy of the modified interf eromet r ic

dispersion measurement, the comparison of the modified interferometric dispersion measurement and the conventional fiber Raman

laser technique has been made, and the results are shown in Fig. 5.

In Fig. 5, the dispersion of A-End ( — ), and B-End ( ) of short

length (1.3 m) are good agreement with that of the fiber Raman

laser technique (- - -) in the long lngth 8.323 (km) of triangularDSF, The measured values of zero-dispersion wavelength (urn),

2 dispersion slope (ps/nm -km) at zero dispersion wavelength in the

modified inter ferometr ic measurement [A-End, B-End] and the conventional fiber Raman laser technique were [1.5000, 1.5008] and 1.5049,

[59.36 x 10" 3

, 59.26 x 10

-3

] and 56.19 x 10~ 3

, respectively.

CONCLUSIONS The high speed and high accurate single-mode fiber

dispersion measuring equipment using modified inerf erometr ic method

was developed. The total measuring time was about 11 (minute) in

the 1.2 um~1.7 urn wavelength regions. The measuring accuracy of

chromatic dispersion of + 0.2 (ps/nm -km) was obtained. Also, it

was found that the dispersion measurement of the inter ferometr ic

method in short length fiber (1.3 m) is equivalent to that of the

fiber Raman laser technique in long length fiber (8.323 km).

ACKNOWLEDGEMENTS The authors wish to thank M. Fujise and M.

Nunokawa of KDD Labs., for their discussions, and M . Ichimura of

SANTEC Corp. for his experimental support, and to Dr. H. Kaji, and

Dr. T. Yamamoto of KDD Labs., for their encouragement.

REFERENCES

[1] M. J. Saunders et al, Tech. Digest, NBS Symp. on Opt. Fiber

Measurements, Boulder, CO, Oct., p. 123, 1984

[2] L. Oksanen et al., ibid, p. 127, Boulder, 1984.

[3] F. M. Sears et al, IEEE, vol. LT-2, No. 2, p. 181, 1984.

[4] Y. Namihira et al,: National Conv. Rec, IECE, No. 1025, 1986 . [5] V.A. Bhagavatula et al., OFC'85, WD-3,p.94, SAN DIEGO, 1985.

[6] L. G. Cohen, IEEE, Vol. LT-3, No. 5, p. 958, 1985.

[7] R. A. Modavis et al., Tech. Digest, NBS Symp. on Opt. Fiber

Measurements, Boulder, CO, Oct., p. 115, 1984

21
Test Single-Mode

F"'ber IR-Visicon (L = 1.3 m ± 5 mm)

Mirror Lens Interferometnc

Ftlter A = 1200~!700nm ^f M°de

Halogen Lamp

InGaAs-APD

<5>-C100W Lens

High Speed

Signal Processing Sub-System

^

Laser = 1 .3 ym

Oiode ^ /-

Single-Mode

Fiber

Corner Cube

Piezo Transducer

Tl t 60~100Hz

d (~10mm) Fig. 2 Appearance of measuring

equipment

.

Fig.l Schematic diagram of measuring equipment

Table 1 Specifications of measuring

equipment

.

Optical source Halogen Lamp

(-100W)

Optical Detector InGaAs-APD

Operating

Wavelength

A = 1200-1 700 nm

per 20 nm, 26 point

(Resolution) S ±0.005(ps/m)

Measuring Accuracy

(Dispersion Ds= + 0.2 (ps/km-nm)

(Zero-Dispersion

Wavelength) Ao^±0.2(nm)

Measuring Time

at one point [Total]

~25 (sec)

[ — 11 minute]

F>' ter Interference Search

A = 1300nm

J= 0 5 10 15 20

Interference position (mm)

Fig. 3 A typical peak search pattern

of interference fringes.

25

E

in

i c

4)E

D.

E

a.

tn Q

(---)

A End 1 Interferometric Method B End J L- 1.300 m)

' Fibre Raman Laser Technique

L- 8.323 km

DSF fl . 3 Triangular

1.00 1.10 1.20 1.30 1.40 1.50 1.60 1.70 1.80 1.90 2.00 Wavelength A Cum]

Fig. 5 Comparison of modified

interferometric dispersion

measurement and fiber

Raman laser technique.

Table 2 Comparison of fitting expressions

1.00 1.10 1.20 1.30 1.40 1.50 1.60 1.70 1.80 1.90 2.00 Wavelength A L«m]

Fig. 4 Delay time and dispersion

characteristics as a function

of wavelength.

Fitting Expressions

Zero-Dispersion Wavelength

I )ispt*rsion-S lope dDU)/<M

x 10' (at An) ps/nm -km

Dispersion

I) Ips/nm - km '

nt A 1.55/im

Fitting Mo.inSipi.ii r ICrror X 10' (us/kin)'

r, (A) 1.5520 65.24 -0.1276 6.134

r2 (A) 1.5525 82.26 -0.2084 0.620

r3 (A) 1.5516 78.31 - 0.1280 0.910

22
CHROMATIC DISPERSION MEASUREMENTS OF SINGLE MODE FIBRES IN THE 1.3 UM AND

1.55 UM WAVELENGTH RANGE BY THE WAVELENGTH TEMPERATURE TUNING TECHNIQUE

Finn Mogensen, Hans Damsgaard and Ole Hansen

NKT Elektronik, Br<Sndbyvesterve j 95, DK-2605 Br0ndby, Denmark

ABSTRACT

Highly accurate (0.5 ps/km for a 4 km length) delay vs. wavelength curves are presented for SM fibres in the 1.3 ym and 1.55 ym windows. is determined

reproducibly with a relative accuracy of 0.02 per cent. An average value

of 1312.8 nm with a standard deviation of 2.8 nm is determined from measurements on production fibres. The dispersion slope S is found to be 0.087

2

2° ps/km-nm with a standard deviation of 0.003 ps/km-nm .

INTRODUCTION

Various experimental techniques for measuring the dispersion properties of

single mode fibres have been reported including Nd : YAG laser system (1) , phase

shift technique using light emitting diodes (LED's) (2), interferometric

method (3) and series of laser diodes (LD's) (4).

In this paper chromatic dispersion measurements are presented for standard SM

fibres and a dispersion shifted fibre with zero dispersion wavelengths around

1.3 ym and 1.55 ym, respectively. The fibres are manufactured by NKT Elektronik using the MCVD process.

EXPERIMENTAL PROCEDURE

The experimental set-up is illustrated in Fig. 1. Two LD's cover the wavelength range around 1.3 ym and a third LD is used in the 1.55 ym wavelength

range. The wavelength tuning is obtained by temperature variation as described previously (5,6).

The delay vs. wavelength curves is measured using the phase shift technique.

The fibre delay is measured relative to a short reference fibre in order to

take into account any (temperature-dependent) delay variations in the LD's.

23
Two main error sources should be considered when using the experimental

configuration shown in Fig. 1. The first is the temperature sensitivity of the

fibre under test (the time delay will change approximately 40 ps/°C-km) and

the second is the frequency stability of the signal generator.

In our experiments the influence of the two above-mentioned error sources was eliminated in one of two ways. Initially the fibre was isolated and temperature stabilized and the measurement time was short (around 10 min.). In this

way accurate and reproducible results were obtained, however, the short

measurement time limited the number of delay points which could be obtained.

Later this was overcome by using a reference LD operating at a constant

wavelength. By the inclusion of the reference LD the temperature stabilization

of the fibre could be avoided and only short term stability was required for

the signal generator.

RESULTS AND DISCUSSION

Fig. 2 shows the relative time delay vs. wavelength measured on an 11 km long

dispersion shifted SM fibre. The results were obtained within approximately 10

min. and no reference LD was applied. Measuring points have been fitted to the

2 -2

usual Sellmeier expression ( a + b A + cA ). The maximum and the typical

deviations between the fitted curve and the measurement points are 0.4 ps/km

and 0.25 ps/km, respectively. The key parameters characterizing the chromatic

dispersion are the zero dispersion wavelength AQ

and the slope S q

of the

dispersion curve at this wavelength. From the fitted curve A = 1545.2 nm and

2 S = 0.061 ps/km-nm were obtained,

o

The reproducibility of the measurement technique was verified by seven repeated measurements on the same 6.5 km length standard 1300 nm optimized

SM fibre. A was determined as 1310.5 nm with a standard deviation of 0.2 nm. o >2 The dispersion slope was found to be S = 0.085 ps/km-nm with a standard

2 ° deviation of 0.001 ps/km-nm .

In order to evaluate the spread in the AQ on tne standard SM fibre from our production, time delay curves were measured for 13 different fibre samples.

The average value of X was found as 1312.8 nm and the standard deviation was

2.8 nm. The average value of S was determined as 0.087 ps/km-nm2 with a ° 2 standard deviation of 0.003 ps/km-nm . The tested samples were from 1-7 km in

length.

24
Finally Fig. 3 shows the delay curve for a 4 km long standard SM fibre

measured using a reference laser. In this case a large number of points (28) were obtained. The maximum and the typical deviations between the fitted curve

and the measurement points are 1.6 ps/km and 0.5 ps/km, respectively. In this

example the chromatic dispersion characteristics were found to be X = 1312.0

2 ° nm and S

q

= 0.090 ps/km-nm .

CONCLUSION

Chromatic dispersion measurements based on the phase shift technique have

been presented. LD's were used as light sources and wavelength variation

was obtained through the temperature tuning effect.

Aq was determined with an accuracy of approximately 0.2 nm in the 1.3 ym

wavelength range yielding a relative accuracy of around 0.02 per cent.

For the standard SM fibre from our production the average value of Xq was determined as 1312.8 nm with a standard deviation of 2.8 nm.

Relative time delay curves were presented both for the standard 1.3 ym SM

fibre and for a dispersion shifted fibre developed for the 1.55 ym wavelength

range. The reproducible and highly accurate (^0.5 ps/km) results confirm that

the temperature tuning technique used here is an efficient alternative to

other methods.

REFERENCES

(1) L.G. Cohen and C. Lin, Applied Optics 16, p. 3136 (1977).

(2) B. Costa, M. Puleo, and E. Vezzoni, Electron. Lett. 19_, p. 1074 (1983).

(3) F.M. Sears, L.G. Cohen and J. Stone, J. Lightwave Tech. LT-2 , p. 181 (1984)

(4) C. Lin, A.R. Tynes, A. Tomita, P.L. Liu and D.L. Philen, Bell Syst.

Tech. J. 62, p. 457 (1983).

(5) T. Miyashita, M. Horiguchi and A. Kawana, Electron. Lett. 13_, p. 227 (1977)

(6) H. Damsgaard, O. Hansen and F. Mogensen, Submitted for publication.

25
TEMP CTRL

LASER

G]

SIG1

GENE

Sz

VALRATOR

FIBRE

COUPLER

APD APD

METER l 2

Fig. 1 Schematic of the experimental set-up.

COMPUTER

PHASE DETECTION

VECTOR VOLTMETER

CHROMA TIC ISPERSION

1520 1530 1540 1550 1550 1570 WAVELENGTH (nm)

CHROMATIC DISPERSION

1270 1280 1290 1300 1310 1320 1330 13-10 WAVELENGTH (nm)

26
MEASUREMENT OF DOPANTS IN OPTICAL FIBERS BY RAMAN SCATTERING

Dan L. Philen

AT&T Bell Laboratories

2000 Northeast Expressway

Norcross, Georgia 30071

INTRODUCTION

This paper describes a scattering technique that provides a new method for determining

dopants in optical fibers. The measurement of the concentration levels of dopants used in

optical fibers is difficult for several reasons. In the region of vibrational spectroscopy beyond

2.5 //m, the fiber is opaque because of the silica lattice vibrations. In the region of electronic

transitions, the fiber is also opaque because of strong UV band-edge absorptions.

Unfortunately, for measurements in the region where the fiber is highly transparent, there are

no strong optical absorptions (that is why the fiber is transparent). Measurements of dopant

levels must then be done by "wet" chemical methods, or by exotic methods such as electron

microprobe (X-ray emission).

A technique that has been used for bulk glass samples [1-4], and preforms [5,6], but has

not found widespread application to optical fibers [7,8], is Raman scattering. Raman scattering

is an inelastic scattering process where a small portion of the scattered light is shifted in

frequency by an amount corresponding to the vibrational and rotational frequencies of the

scattering medium. The Raman scattered light shifted to frequencies higher than the excitation

is the anti-Stokes component, and those shifted to lower frequencies than the excitation is the

Stokes component. The strength of the scattered light is given by m = a F ; where a is the

polarizability of the molecule, and F is the electric field strength.

Raman scattering occurs at any excitation wavelength, so in principle any

monochromatic excitation source can be used, but. the intensity of the scattering increases with

decreasing excitation wavelength, and the scattered information is contained within a smaller

wavelength region. The scattered light is shifted in energy from the excitation line by an

amount proportional to the vibrational frequency involved. Thus, a 1500 cm'1 vibration would

appear at 425.5 nm when excited by a 400 nm line (25.5 nm separation), but at 659.3 nm when

?1
excited by a 600 nm line (59.3 nm separation). There is a compromise between the intensity of

the scattering and the resolution from the excitation line that can be obtained with

monochromators.

Since glasses are not good scatterers, sample size has been a problem when dealing with

bulk samples. It is difficult to get enough glass material to interact with the excitation light and

then efficiently collect the scattered light by a monochromator. Thus, complicated collection

optics are generally required for studies on bulk samples. Optical fibers represent a solution to

this difficulty because focusing the light into the small volume of the fiber core increases the

electric field density, and scattering intensity is directly proportional to field strength. The

excitation light is confined to the fiber core, so long interaction lengths are possible. Since the

scattered light can be analyzed by forward scattering (as opposed to perpendicular scattering)

sample sizes from a meter to several hundred meters are possible. Complicated light collection

systems are not needed because the output end of the fiber can be placed at the entrance slit of

the monochromator and all the forward scattered light detected.

EXPERIMENTAL

A simple Raman scattering experiment is shown in Figure 1. Monochromatic light from a

laser source is launched into the test fiber, and the scattered light exiting the fiber is directed

into a double monochromator. The monochromator is scanned, starting near the excitation line,

toward longer wavelengths. The vibrational bands of interest are observed displaced from the

excitation line by the energy of the vibration. A photomultiplier is used as the detector and the

output plotted on a strip chart recorder using phase sensitive electronics.

RESULTS

In Figure 2 the observed Raman scattering bands for several experimentally determined

vibrations are shown. The vibrational assignments are also shown in the accompanying table.

Of particular interest is the vibration at 1320 cm'1

, assigned to the P=0 stretching vibration.

The other observed transitions are broadened by the coupling of the vibration to the silica

network. However, the P=0 vibration is much sharper than the other observed vibrations

because the P=0 bond is a terminating link in the network. The oxygen is not a bridging

28
oxygen as in the Si-O-Si bridge, and thus the P=0 vibration is not damped by the rest of the

silica network, and appears as a sharp transition. Also, this vibration appears in a region where

there is little other interference from other dopants. We have used the uniqueness of this line

for quantitative analysis of phosphorus levels in test fibers down to 0.05 mole percent, and think

that quantitative determinations could be made to significantly lower levels.

One advantage to this type of analysis is that the silica can be used as an internal

calibration. Since silica is in such a great excess compared to the other dopants present, the

silica scattering peaks at 1060 cm'1 or 1600 cm'1 can be used as internal standards and the

unknown peaks compared to them. For the results presented here, the ratio of the scattering

peaks was determined for a calibration sample analyzed by electron microprobe. Subsequent

samples were then easily computed by determining the peak ratios and comparing them to the

calibration sample. The sample peak ratios are independent of the sample size since we are

comparing to a known silica reference within each sample. A larger sample (longer fiber length)

does give increased signal to noise however, and data will be presented using a 5 mW He-Ne

laser and 60 meters of sample.

For most fibers, 10 meters of sample is enough, and a spectral scan requires about twenty

minutes. Qualitative measurements are easily made for common dopants such as germania,

phosphorus, fluorine, and boron, with quantitative determinations being only slightly more

difficult. Quantitative measurements need suitable standards and peak-area determinations.

CONCLUSIONS

Raman scattering is ideally suited to the study of optical fibers because (l) the excitation

light is not confined to any particular region of the spectrum (2) sample size is greatly simplified

since the light is confined to the fiber core, and (3) no complicated collection optics are needed.

In addition, calibration samples need only be measured once since the ratio of peaks within the

test sample may be used as an internal reference. Quantitative determinations of phosphorus

levels have been made down to 0.05 ± 0.02 mole percent using this technique. The

determination of other dopants like fluorine, germania, and boron is also equally possible.

29
REFERENCES

1. M. C. Tobin and T. Baak, "Raman Spectra of Some Low-Expansion Glasses," J. Opt. Soc. Am, 58, 1459 (1968).

2. M. C. Tobin and T. Baak, "Raman Spectra of Vitreous B2 03 and of Some B2 03 ~Al 2 Os Glasses," J. Opt. Soc. Am. 60, 368 (1970).

3. F. L. Galeener, J. C. Mikkelsen, R. H. Geils, and W. J. Mosby, "The Relative Raman Cross

Sections of Vitreous Si02, Ge02 , B2 Oz

, and P2 Ob

," Appl. Phys. Lett., 32, 34 (1978).

4. N. Shibata, M. Horiguchi, and T. Edahiro, "Raman Spectra of Binary High Silica Glasses and

Fibers Containing Ge02 , P2 Oh , and 52 03

," J. Non-Cryst. Solids, 45, 115 (1981).

5. W. A. Sproson, K. B. Lyons, and J. W. Fleming, "Nondestructive Concentration Profiling of Fiber Optic Preforms by Analysis of Raman Spectra," J. Non-Cryst. Solids, 45, 69 (1981).

6. P. Dumas, J. Corset, W. Carvalho, Y. Levy, and Y. Neuman, "Fluorine Doped Vitreous Silica Analysis of Fiber Optic Preforms by Vibrational Spectroscopy," J. Non-Cryst. Solids, 47, 239

(1982).

7. G. W. Bibby and J. N. Ross, "Raman Spectra of Hydrogen Treated Optical Fibers,"

Electron. Lett., 20, 182 (1984).

8. K. Noguchi, Y. Murakami, N. Uesugi, and K. Ishihara, "Raman Scattering Spectrum

Analysis for Fluorine-Doped Silica Optical Fibers," Appl. Phys. Lett., 44, 491 (1984).

. Si-O-Si

-J

Recorder

750 1000 1250 1500 2000 FREQUENCY (cm-')

Experimental Set-Up FIGURE 2

FIGURE 1

30
UTILIZING CLADDING MODE PROPAGATION TO DETERMINE THE REFRACTIVE

INDEX PROFILE OF AN OPTICAL FIBER

W.H. Hatton, E.L. Buckland and M. Nishimura

Sumitomo Electric Research Triangle, Inc.

Research Triangle Park, NC 27709, USA

1. Introduction

A number of measurement techniques have been developed [1-2]

in order to ascertain the refractive index profile of an optical

fiber. For the most part, these measurement methods are

associated with numerous drawback ( i.e. elaborate sample

preparation, correction factors) . The cladding mode near-field

(CNF) technique has been developed in order to reduce the overall

complexity of index profiling measurements. The CNF technique

gives a direct measure of the refractive index profile, similar

to the refractive near-field method[3-4], however, sophisticated

optical and mechanical equipment are not necessary.

2

.

Principle of the CNF Technique

By surrounding an un-coated optical fiber with a material

whose refractive index is less than the cladding ( for example

air) light will not only propagate in the core, but also through

the cladding region ( i.e. cladding modes) . Consider the

propagation of only cladding modes ( no modes bound by the core ) as shown in Figure 1. Utilizing Snell's law, light ray

trajectories at the output end of the test fiber can be

31
described by

No cos 0,= N(r) cos02 (1)

N(r) sin02 = sin 0 3 (2)

where N(r) is the local refractive index along the exit plane of

the fiber, and No is the refractive index of the cladding. By

combining equations (1) and (2) , the angle at which light exits

the fiber, 63 , can be uniquely related to a propagation angle of

incidence 0i

.

2 2 2 2

N(r) - No = -No sin Q\ + sin $3 (3)

If uniform excitation of cladding modes exist then, from

equation (3) , the magnitude of N(r) at any radial position along

the endface of the fiber determines the relationship between 9»

and 03 . Assuming that No and sinQj are fixed, any variation in

the fiber's refractive index profile will induce an equivalent

change in sin0i. By examining the outputed intensity distribution

with a detection system whose numerical aperture is limited

(fixed at sin ©3 )

/ total power at any arbitrary point along the

detected pattern will be directly related to 0i

3 . Description of CNF Measurement System

A schematic drawing of the measurement system is shown in

Figure 2 . Light from an intense optical source is focused along

the glass/coating material interface of the test fiber. A small

fraction of the power scattered at this interface will be coupled

into the cladding region where for short lengths of un-coated

32
fiber, only cladding modes will be excited. The power outputed

from the test fiber is projected onto a vidicon TV camera system

where the total intensity profile is analyzed. The numerical

aperture of the detection system is limited by an aperture,

Al, placed directly behind lens L2 . This position was chosen

for convenience and is equivalent to positioning the aperture in

front of the lens.

Preliminary laboratory results from the CNF technique have

shown to be accurate when compared to data from other measurement

methods. Three dimensional raster scans, as shown in Figure 3

for a polarization maintaining single mode fiber, are also easily

obtained.

4 . Conclusion

A new method for accurately measuring the refractive index

profiles of optical fibers has been developed. The cladding mode

near-field technique has been shown to give a direct measure of

the index profile without the requirement of elaborate test

equipment, therefore, overall complexity of this type of

measurement can be reduced.

References

[1] D. Marcuse et al.,Proc. of IEEE, 68, No. 6, pp.666 - 688

(1980)

[2] W.J. Stewart, IEEE Trans. Microwave Theory Tech., MTT-30,

pp. 1439-1454 (1982)

[3] W.J. Stewart, in Technical Digest of IOOC'77, paper C2-2

(1977)

[4] K.I. White, Opt. Quant. Elec. , 11, pp 185-196 (1979)

33
34
DETERMINATION OF REFRACTIVE -INDEX DISTRIBUTION

FROM NEAR-FIELD MEASUREMENTS IN POLARIZATION MAINTAINING FIBRES

G. Coppa, P. Di Vita, M. Potenza

CSELT - Via G. Reiss Romoli, 274 - 10148 Torino (Italy)

Introduction

Polarization maintaining single-mode fibres are gaining increasing interest

both in the field of sensors and in the field of communications. Their design

is often based on azimuthally dependent refractive-index distributions: then

these fibres can be characterized obtaining their two-dimensional

refractive-index distribution.

In this contribution we propose an alternative method to the classical

Refracted Near-Field technique [1], based on a numerical elaboration of the

(Transmitted) Near-Field (TNF) intensity pattern of the fibre. This method has

been already employed for circularly symmetric single-mode fibres and its

effectiveness for one-dimensional index profiles has been proved [2], However

its extrapolation to two-dimensional index distribution measurements is not

trivial, since the acquisition of a matrix of at least 50x50 TNF-points (with

S/N > 30 dB) is needed. This could be accomplished by a TV camera and a

digitizing system, but the required performances are nearly at the limits of

the available instrumentation. Otherwise a raster scanning of the measurement

matrix could be performed by a photodiode. This leads to a considerably

greater measurement time with respect to the one-dimensional case but it does,

not represent a particularly serious problem since available stabilized optical

sources and precision motorized displacement scanning stages are required.

35
Moreover, Che signal processing needs smoothing and differentiation algorithms

which are substantially different owing to the two variables involved.

Nevertheless the method is very convenient, since it requires a simple TNF

set-up which is usually employed for the measurement of different parameters

(e.g. mode field diameter or geometrical characteristics).

Theory

The TNF-data processing is based on the fact that in single-mode fibres the

intensity I(x,y) is proportional to the square of a component of the

electromagnetic field E(x,y) which satisfies the following scalar wave

equation

( & +il)E(x,y) + {[2Ttn(x,y)/^J 2

'- |^}E(x,y) = 0 (1)

n(x,y) being the refractive index distribution, % the operating wavelength and

[b the mode propagation constant. Then it is possible to obtain the quantity

{ [2 31 n(x,y )/% ] - (s } from Eq. (1) directly, in terms of the measured field

amplitudes. Since second derivatives have to be performed numerically on the

measured data, these should be very accurate and a suitable smoothing procedure

could be necessary. Absolute refractive- index differences and fibre NA's can

be derived.

Experimental results

The experimental apparatus is a standard TNF intensity scanning set-up with a

dynamical range around 40 dB. The fibre under test is excited by an LEU, its

output end is magnified by means of a 0.65 NA objective and an x-y scan is

performed by a moving photodiode in the image plane. A desk-top computer

drives the displacement stages, acquires the TNF-data and processes them.

36
Figs. 1 to 3 show the results of the present method in some meaningful cases.

In each figure the calculated refractive index distribution (a) and the

TNF-pattern (b) for each fibre are reported. Fig. 1 concerns a circularly

symmetric single mode fibre ( % =1.3 ym ): the accuracy of the circular

geometry confirms the good quality of the measurement procedure. Two fibres

made in our laboratories are considered in Figs. 2 and 3 ( % = 1.5 ^im ). The

first one has an elliptical core with a remarkable dip in the centre that

causes the field to break up into two peaks. Note how the depressed regions in

the cladding are accurately evidenced by our algorithm. Finally Fig. 3 shows

the characteristics of a square core fibre: though the structure is very

complicated, an index distribution consistent with the original fibre design

has been derived in this case too.

Conclusions

A simple technique for the determination of two-dimensional refractive-index

distributions in optical fibres has been proposed. It is based on a suitable

elaboration of near-field measurement data. Experimental results have shown

the accuracy and the reliability of the technique. This method can represent

an useful tool of investigation for polarization maintaining optical fibres

and, in general, for any non-circularly symmetric lightguides.

References

1. W. J. Stewart, IOOC'77, Tokyo , Japan (1977), Tech. Digest, pp. 395 - 398.

2. G. Coppa et al.,IOOC'83, Tokyo, Japan (1983), Tech. Digest, pp. 38-39.

37
s /a s /a x C///11J

Fig. 1 Refractive index distribution (a) and two-dimensional near-field

pattern (b) of a circularly symmetric fibre. A is the maximum NA.

4=0.20

Fig. 2 The same as Fig. 1 but for an elliptical core fibre.

4=0.13

Fig. 3 The same as Fig. 1 but for a square core fibre.

38
Prism coupling technique: A novel method for measurement of

propagation constant and beat length in single mode fibers

K.Thyagarajan, M.R. Shenoy and M.R. Ramadas

Physics Department

Indian Institute of Technology

New Delhi 110016, INDIA

The prism-film coupling technique in integrated optics

is a well-known technique for studying the mode spectra of

planar optical waveguides, A prism whose refractive index is

greater than or equal to the maximum refractive index of the

waveguide is positioned close to the waveguiding structure;

light incident on the base of the prism is coupled into the

waveguide through the evanescent field in the intermediate layer.

In a normal cladded fiber the modal field is not accessible

from outside the fiber due to the large cladding thickness.

However, if one can remove the cladding so as to be in the

proximity of the core, it would be possible to couple light into

(or out of ) the fiber using this arrangement. Such a technique

has been used earlier [1] to selectively excite particular

groups of modes in a multimode fiber.

In this paper, we report preliminary investigations on

the application of the prism-coupling technique to the measurement of the mode spectra of conventional and birefringent single

mode fibers (SMF). Experimental results on the measurement of

propagation constant in normal 3 IF, and the application of this

technique to the measurement of beatlength on an elliptic core

SMF, is presented.

In order to use the prism coupling technique, the cladding of the fibers were selectively removed using the standard

polishing technique. Two single mode fibers, one from York

39
Technology (SMoOO) and the other from I IT (TloOl), having a cut

off wavelength «~ 600 nm were used in the experiment. The

refractive index of the prism used was 1.5780. Light from a He-Ne laser was coupled into the fiber from one end and was

decoupled using the prism placed on the polished surface. The

propagation constant of the fiber mode can be determined by

measuring the angle 0 at which the mode comes out of the

prism (see Fig.l). The measured values of the propagation

constants of the two fibers together with the calculated values

using the fiber parameters given by the manufacturer and measured

in the laboratory (for the ITT fiber) are given in table I. As

can be seen the agreement is remarkably good; the slight difference can be attributed to the possible deviation of the prism

lower surface from the plane tangent to the fiber's curvature,

and also to the perturbation caused by the polishing operation.

In the present experiment, however, the perturbation due to

polishing is estimated to be small due to the observation that

inspite of using an index matching liquid between the prism and

the polished surface, complete decoupling of the mode was not

possible which implies that a sufficient thickness of the cladding was still left.

In order to show the applicability of the technique to

the measurement of beat length in birefringent fibers, we

polished an elliptical core fiber (No.850417A- from Andrew

Corporation) with its major axis perpendicular to the polished

surface. Light from a He-Ne laser with the plane of polarization at 45° to the eigen polarization axes of the fiber was

coupled into the fiber; the output consisted of two closely

spaced ;n-lines separated by an angle of ^4*, and each one

40
was polarized, one with its plane of polarization parallel to

and the other perpendicular to the polished surface. From the

output coupling angles one can determine the difference in

effective indices { An ) of the modes which would give us the

beat length according to the relation 1^ = ^/ao . The

beat length of the same fiber was also measured independentlyusing the Rayleigh scattering technique. The results of both

these measurements are given in table II. As can again be

seen, the agreement is very good.

The proposed novel technique should find applications

in the measurement of the cut off wavelength in single mode

fibers, the fiber dispersion characteristics, and can also be

used to estimate the effect of polishing on the birefringence

characteristics of single mode fibers 2 J • The authors thank Dr. rt.B. Dyott of the Andrew Corporation for providing samples of elliptic core fibers, and

Professor A.K. Ghatak for constant encouragement.

References :

1 J.E. Midwinter, Optical Fibers for Transmission , John

Wiley & Sons, New York (1979), p. 208.

2 H.H. Stolen, ' Polishing-Induced birefringence in Singlemode fibers,' Applied Optics 25, 344 (1986).

Table I

Fiber type

Measured Calculated

from manuf acturers

data

Calculated from

data obtained by

RNF technique

ITT

York

Technology

1.4592 + 0.0002

1.4590 + 0.0002

1.45902 1.45927

41
Table II

Fiber Beat length measured using

Prism coupling Rayleigh Scattering

Elliptic Core 1.52 +0.08 mm 1.5 + 0.C1 mm

Fig.l

42
CHARACTERIZATION OF OPTICAL FIBER SPLICES IN THE FIELD

Rein hard Engel

Siemens AG,

Munich, W Germany

Everett McNair

Siecor Optical Cable

Hickory, NC 28603

I . INTRODUCTION

The objective of field splicing operations is to attain a

permanent jointing which exhibits high strength and low loss. The

generally recognized steps which must be accomplished to achieve this

objective are; (1) precision alignment of the waveguides, (2) joining

the waveguides, (3) permanent retention of the fiber's alignment and (4)

an evaluation of the quality of the splice. This paper addresses field

splice loss evaluation by reviewing the factors which contribute to

splice loss and discussing some of the splice loss measurement methods

that are currently being used in the field.

II. SPLICE LOSS CHARACTERIZATION

Splice loss arises from both intrinsic and extrinsic factors.

Loss attributed to intrinsic factors results from fiber-to-fiber

parameter variations which occur during the manufacturing process. The

degree to which intrinsic fiber parameters contribute to insertion loss

at a splice can not be controlled or practically determined in the

typical field environment.

Extrinsic contributors to splice loss are factors which can be

directly influenced by the craftsperson or splicing technique. Lateral

or axial misalignment of the fiber cores is the most significant

extrinsic factor.

During the the splicing process the intrinsic and extrinsic losses

are not additive. Instead they combine to yield a total loss which is

less than their algebraic sum [1].

43
III. Splice Loss Evaluation in the Field

Transmi ssion Method

.

Fig. 1 illustrates a laboratory procedure

for evaluating splice loss between dissimilar fibers. Before making the

splice, the power, PI, transmitted through the first fiber is measured.

After completing the splice a second power measurement, P2 , is obtained

in the immediate vicinity of the splice. The splice loss is then

determined from

Loss = -10 logP2/Pl dB (1) .

This procedure is adaptable to the field environment but for multimode

fibers, the location and the nature of the power source relative to the

splice as well as the length (L2) of the fiber which follows the splice

can affect the results obtained from transmission measurements [2]. For

accurate splice loss measurements, the respective attenuation of the

individual fiber lengths must be accounted for. In the field, this

process requires a minimum crew size of three people. The uncertainties

associated with cable losses can be eliminated by an ingenious but time

consuming procedure known as prespl icing [3], with a reported accuracy

of +/- 0.02dB.

Local Dectection

.

The use of local scattered light detection [4]

reduces the overall splicing process to a two person operation and

improves the speed with which two singlemode fibers can be aligned.

Fig. 2 [4] provides a comparison of the alignment sensitivities achieved

by far end transmission detection and local scattering detection

methods. Splice loss is derived from a correlation between the amount

of light coupled into the far end of one of the fibers and the amount of

light that is locally detected at the splice point.

Optical Time Domain Refl ectometry

.

Although the OTDR's accuracy

is limited, splice loss can be easily evaluated by accessing only one

44
end of the jointed fibers. The OTDR measurement errors can be reduced,

at the expense of productivity, by evaluating the splice from both

directions and taking the mean of the results. Table 1 [5] shows the

variances which can result from OTDR uni -directional measurements. An

average directional splice loss difference of 0.20 dB was obtained from

bi-directional measurements of 11 single mode splices. The maximum

variance in measurements taken from opposite directions is 0.50 dB. The

average of each series of uni -directional measurements compares

favorably with the bi-directional average, illustrating that the

uni -directional inaccuracies average out as the number of splice

measurements increases.

Local Injection and Detection

.

The most reliable method of

aligning fiber cores is the transmission method, because this measuring

method duplicates the operating state of the fibers [6]. A prerequisite

for its cost effective field application is the launching of light into

the fiber core before the splice point and the extraction after the

splice point. Application of bending couplers allows the injection of

measuring light through the fiber cladding and coating into the fiber

core directly at the splicer unit [7]. A local injection and detection

system (Fig. 3) simplifies the splicing process to a one-person onelocation operation

.

Some of the factors that affect measurement accuracy for single

mode fibers are the measuring wavelength versus the cut-off wavelength

for the fiber, amount of mode scrambling between the injection device

and the splice, and the existence of cladding modes.

Given constant injection conditions, the extraction device detects

the optical power level ,Pa, after the fiber ends have been fused. Using

Pa and the ideal optical power level ,Pi , the splice loss is calculated

45
as Loss = -10 1 ogPa/Pi dB (2)

.

Pi is the level which could have been detected by the extraction device,

if the splice had no loss. Therefore, the most accurate determination

of the ideal level Pi, is a precondition for the calculation of splice

loss

.

One possible solution is to produce a reference splice by applying

a drop of index matching fluid between the end faces of the fibers.

This almost completely suppresses reflections caused by the air space

between the fibers and perturbations due to angled end faces. After

alignment in the lateral and axial directions, the power level at the

detector corresponds so closely to the ideal level, that with identical

single mode fibers the calculated splice loss has a typical margin of

error of +/- O.OldB [8].

When index matching fluid is not used, the opposing end faces form

an optical resonator which reduces the optical transmission between the

fiber ends by up to 0.6 dB, depending on the distance between the end

faces. If the distance is greater than the coherence length of the

measuring light, the interferences disappear and light transmission is

only diminished by Fresnel reflection (0.3 dB), by the quality of the

end faces and the gap between them (Fig. 4). If applied to identical

single mode fibers, calculation of the ideal level according to the

"gap" method for determination of the splice loss leads to a margin of

error of +/- 0.08 dB [9]. As in the case of the "wet" method, here too

attenuation due to optical mismatching is not taken into consi sderation

.

Much more accurate calculation of the ideal level, which also

would take intrinsic losses into consideration, is expected from another

"dry" method. With it, the interference response of the measuring light

between the fiber end faces is used for calculation of the ideal level .

46
A maximum error of +/-0.015 dB is expected.

Direct Core Monitoring

The direct core monitoring [10], [11]

splicing technique takes the approach of obviating splice loss measurements by addressing the accuracy with which fibers can be aligned solely

by observation of the fiber cores. The estimation is based upon an

analysis of the accuracy with which the fiber cores can be observed and

manipulated to minimize alignment errors. Reportedly [10], a spliceloss standard of .2 dB can be guaranteed with 99 per-cent reliability

when the observation detection accuracy is kept within 0.25 urn.

IV. SUMMARY

Field procedures for evaluating the quality of a splice must enhance the craftsperson 1 s productivity as well as deliver the highest

practical measurement accuracy. Ideally, the selected splicing technique is designed so that all four steps involved in the splicing operation can be performed by one individual located at the splice point.

Local injection and detection systems offer this convenience and studies

are on-going to further enhance the benefits of this technique.

REFERENCES

[1] CM. Miller, "Optical Fiber Cables and Splices," IEEE J. on

Sel . Areas in Comm

, Vol. SAC-1, No. 3, Apr. 1983; p . "5T6

[2] R.B. Kummer , A.F. Judy, and A.H. Cherin, "Field and Laboratory

Transmission and OTDR Splice Loss Measurements of Multimode

Optical Fibers," Tech. Digest-Sym. Opt. Fiber Meas. , 1982; pp.

109-121

[3] Y. Kato, S. Seikai , and T. Tanifuji, "Arc-Fusion Splicing of

Single-Mode Fibers: An Apparatus with an Automatic Core-Axis

Alignment Mechanism and Its Field Trial Results," IEEE J. LW.

Tech. , Vol. LT-2, No. 4, Aug. 1984

[4] CM. Miller, "Local detection device for single-mode fiber

splicing," Proceedings of OFC '82, paper THAA2

[5] J. Lidh, "Splice losses: How measurements can mislead," Lig htwave, The Jour, of Fiber Opts., June 1986 ; p . 47

[6] T. Pavlopoulos, A. Lewis, M. McLandrich, "Simple method for

permanently connecting single-mode fibers," Appl ied Optics , Vol . 16, No. 6, 1977; p. 1466

47
[7] C. DeBlock, P. Matthysse, "Core Alignment Procedure for SingleMode-Fibre Jointing," El ec . Ltrs. , Vol . 20 Mo. 3, 1984; p. 109

[8] Y.Kato, T. Tanifuji, M . Tokuda, N. Uchida, "New Optical Monitoring

for Arc-Fusion Splice of Single-Mode Fibbres and High-Precision

Estimation of Splice Loss," El ec . Letters. , Vol. 18, No. 22,

1982; p. 972

[9] Y. Kato, S. Seikai, T. Tanifuji, "Arc-Fusion Splicing of

Single-Mode Fibers: An Apparatus with an Automatic Core-Axis

Alignment and its Field Trial Results," Jour . of LW. Techn . , Vol . LT-2, No. 4, 1984; p. 442

[10] 0. Kawata, K . Hoshino, Y. Miyajima, M. Ohnishi, and K . Ishihara,

"A Splicing and Inspection Technique for Single-Mode Fibers Using

Direct Core Monitoring," IEEE J. LW. Tech. , Vol . LT-2, No . 2

,

Apr. 1984.

[11] 0. Kawata, K . Hoshino, and K. Ishihara, "Low-Loss Single-ModeFibre Splicing Technique Using Core Direct Monitoring," El ec

.

Letters, Vol. 19, Nov. 24, 1983.

Source

Optical Power Meter

Splice Saurte r ^T— Optica) Power Meter

Splice Lois - -10 log -==- dB

Laboratanj splice Lass Measurement

for Oistimiler Fibers

Sensitivitg Comparison

(Single Mode Fiber)

Traniuerse Offset luml

Figure 2

Ftialadioa'* -

lool ln|.cllon mnd Dlltillon »«•••" 'nncipH

j nquf i

|

VARIATIONS IN 0TDR MEASUREMENTS

SPLICE

NO. MEASUREMENTDIRECTION

ABSOLUTE DIFFERENCE BIDIRECTIONAL

1 .31 .24 .07 .28

2 -.22 .28 .50 .03

3 .10 .04 .06 .07

4 .19 -.06 .25 .07

S .00 .28 .28 .14

6 .17 .15 .02 .16

7 .10 .01 .09 .06 8 .27 -.15 .42 .06

9 .07 .08 .01 .08

10 .15 -.11 .26 .02 11 .00 .30 .30 .15

AVG .104 .096 .205 .100

TABLE 1 VARIATIONS IN OTDR MEASUREMENTS

(From R»f. CS])

figure 4

48
CONSISTENCY OF MODE FIELD DIAMETER DEFINITIONS

FOR FIBERS WITH GAUSSIAN AND NON-GAUSSIAN FIELD PROFILES

IN DIFFERENT MEASUREMENT DOMAINS

by W. T. Anderson and J. P. Kilmer

Bell Communications Research

1. Introduction

The Mode Field Diameters (MFDs) of two single- mode fibers can be used to predict the joint loss [l]. Controlling the MFD in manufacture will limit the intrinsic splice loss between two single-mode fibers from one supplier, and the differences between the MFDs of fibers from different manufacturers is a measure of the degree of compatibility between them. However, for the MFD to be of practical value as a predictor of fiber compatibility, the MFD measurements made by suppliers and users must be

consistent for all fiber designs. Four measurement methods are commonly used: far-field scanning [2], measuring transmission through an offset joint [3], measuring transmission through apertures in the farfield [4,5], and far-field knife-edge scanning [6]. After years of debates in national and international

standards organizations, none of these methods is universally accepted as a reference test method, nor is any method more or less fundamentally correct than the others. Therefore, consistency among the

methods is especially important.

The MFD is computed from data generated by any of these measurement methods by applying a

suitable definition. In this paper, we consider two different definitions. For the first generation of single-mode fibers, whose index of refraction profiles are nearly step-index, the field distributions are well

approximated by gaussian functions [7]. It has been shown that, for the first generation of single-mode

fibers, these gaussian definitions give consistent results among several of the measurement methods [8,9]. An alternative definition of the Mode Field Diameter is based upon the second moment of the far fields

[10,11,12]. This second moment definition, which is often called the "Petermann 2" definition, is independent of the shape of the field distribution and, in principle, can be used to predict the joint loss between any two single-mode fibers of similar design. However, the applicability and consistency of this

definition to the commonly used measurement methods has not been established, although a theoretical

study [13] has shown that this definition is more consistent than the gaussian definition for fibers with

non-step-index profiles.

The measurement methods are related to a good approximation by a number of transforms, as shown in Figure 1. Because of these transform relationships, theoretically or experimentally derived fields for any

domain may be transformed into the other domains, and the MFDs which result from applying the two

definitions in all five domains may be compared. In this manner, the degree of consistency among the

measurement methods can be quantified, and a rational basis for the selection of a definition for the

MFD can be established.

In section 2, MFD consistency is tested for a perfect step-index fiber using theoretically computed far

fields. These results are verified using measured far fields for two commercially available nearly stepindex fibers. In section 3, the MFD consistency is tested using the measured fields for. a commercially

available non-step-index dispersion shifted fiber. Finally, in section 4, conclusions are drawn.

49
2. Step-Index Fibers

The gaussian definition is applied to the four measurement methods as follows: far field- the overlap

integral, offset joint- the 1/e point, variable aperture- an integrated gaussian fit, and knife edge- an

error function fit. For the second moment definition, moments can be used in all but the offset joint

measurement, where the central curvature is needed.

The far fields of a step-index fiber can be expressed in closed form, so this is a logical starting point for the theoretical study. With a single integration or transformation of the far fields, the offset joint,

variable aperture, or knife edge measurement can be simulated, and the consistency of the definitions can be tested by applying the definitions in each of the domains and comparing the resulting MFDs.

Similarly, measured far field data for two approximately step-index fibers can be transformed into the

other three domains and MFDs can be determined in an identical fashion. One of the two fibers is a depressed-cladding design manufactured using the inside process, and the other is a matched-cladding

design manufactured using the outside process. When the second moment definition is applied to a set of transformed far field data, a consistent MFD value (to within 0.2%) is calculated in each domain.

This is expected due to the portability of the moment definition. This moment MFD value is used as a gauge to compare the MFDs which result from applying the gaussian definition to the theoretical and

experimental data for the two fibers at the two measured wavelengths in the four measurement

domains. This is shown in figure 2. The gaussian MFDs are fairly consistent at 1300nm, but are less consistent at 1540nm where the field distributions are evidently somewhat less gaussian. Since the usual

tolerance on the MFD is ±10%, then differences of less than one or two percent are desirable. The

theory and experiment are in good agreement considering that the fibers which were measured differed from an idealized step index profile.

3. Non-Step-Index Fibers

Far fields for a dispersion-shifted fiber [14] were measured. After transforming this data to the other

three domains, MFDs were computed using both the gaussian and the second moment definitions. Results are shown in figure 3. The gaussian definition gives inconsistencies as large as 15%, while the

second moment definition is consistent to within 0.2%. To verify that the second moment definition is a

better predictor of splice loss, the autocorrelation function for this fiber is plotted in figure 4. Also

plotted are the splice loss estimates resulting from both the second moment MFDs and the gaussian

MFDs. Clearly, the second moment definition gives a superior estimate of splice loss.

4. Conclusions

For step-index fibers at 1300nm, either the gaussian or the moment MFD definition leads to consistency

among the five measurement methods. However, for fibers with non-gaussian fields, the gaussian

definition leads to inconsistencies as large as 15% between measurement methods. Furthermore, the

gaussian definition does not lead to accurate predictions of splice loss. The next step in this study is a

direct measurement comparison among the various methods to be certain that actual implementations

yield the levels of consistency which the transforms have shown to be possible. Such a direct comparison will be presented in a forthcoming paper.

50
REFERENCES

[I] D. Marcuse, BSTJ, Vol. 56, No. 5, p. 703, 1977.

[2] K. Hotate and T. Okoshi, Applied Optics, Vol. 18, No. 19, p. 3265, 1979.

[3] J. Streckert, Optics Letters, Vol. 5, No. 12, p. 505, 1980.

[4] E. Nicolaisen and P. Danielsen, Electronics Letters, Vol. 19, No. 1, p. 27, 1983.

[5] D. K. Smith and R. A. Westwig, Technical Digest, OFC 85, Paper TuB2, 1985.

[6] P. Samson, Optics Letters, Vol. 10, No. 6, p. 306, 1985.

[7] D. Marcuse, J Opt Soc Am, Vol. 68, No. 1, p. 103, 1978.

[8] W. T. Anderson and D. L. Philen, IEEE JLT, Vol. 1, No. 1, p. 20, 1983.

[9] D. L. Frantzen and R. Srivastava, IEEE JLT, Vol. LT-3, No. 5, p. 1073, 1985.

[10] K. Petermann, Electronics Letters, Vol. 19, No. 18, p. 712, 1983.

[II] C. Pask, Electronics Letters, Vol. 20, p. 144, 1984.

[12] W. T. Anderson, IEEE JLT, Vol. LT-2, No. 2, p. 191, 1984.

[13] J. C. Auge and L. B. Jeunhomme, Technical Digest, SPIE Tech Symp, August, 1985.

[14] T. D. Croft, J. E. Ritter, and V. A. Bhagavatula, Technical Digest, OFC 85, Paper WD2, 1985.

Figure 4

OFFSET (MICROMETERS51
Matched Oao Sltp Inflei Fiber = 10um and _ = 0 3 at !300nm

Depreaaed Clad Step index Fiber d=B 3um and £ — 0 37 al 1300nm

J.

Figure 2

Theoretically Calculated Fartieid Data

Experimental Measured Fartield Data

Matched Clad Step index Fiber lOum and c TM 0 3 at I540nm

mil

15 <4> j_

Depresaed Clad Step Index Fiber

B 3um and o 0 37 al lS40nm

Far Fwld

Oauaaian

Knife Edga

Error Function Tranaveme

Oriaei

1 • wttfth

Varaible Aperture Geuaaian

52
A NEW TECHNIQUE FOR MODE FIELD DIAMETER DETERMINATION

BY TRANSVERSE OFFSET MEASUREMENTS

M. Calzavara, G. Coppa, P. Di Vita

CSELT - Via G. Reiss Romoli, 274 - 10148 Torino (Italy)

Introduction. The problem of standardization of the mode field diameter (MFD)

of single mode fibres is still debated. Different definitions of MFD are

available with different meanings [1]. Among these the MFD d defined as the

inverse of the r.m.s. width of the far-field distribution is finding an

increasing interest, d can be evaluated in principle from Car-field, from

near-field and from transverse offset (TO) measurements. At present the last

measurement technique seems the most critical. In fact calling P(u) the

coupled power between the two fibre sections with a pure transverse offset u, d

is given by [2]:

(1) d=2[-2P(0)/P"(0)] </5'

The most convenient way to calculate d consists in fitting the measured P(u) by

a suitable function for u<<d and applying Eq. 1 to the fitting function. This

requires a lot of experimental points in the fit region. Although many

practical problems of the TO technique have been overcome by the self-imaging

version of the method [3] the need of submicron displacements still could

represent a weakness of this technique with respect to the others. In this

contribution we present some substantial improvements to the self-imaging TO

technique that overcomes this weakness. The fitting procedure of the TO curve

is also discussed and an effective algorithm is suggested.

Measurement technique . The chosen TO set-up (Fig. 1) is the self-imaging

version further improved by a new optical component that simplifies the fibre

53
positioning procedure and improves optical resolution, and by introducing h

fast scanning high-resolution measurement of TO curves.

The new optical component L (Fig. 1 ) is both reflective and refractive. It is

obtained from a positive aplanatic meniscus lens, known also as Luboshez lens

[4], by means of a partially reflecting metallic layer deposited on the concave

surface. The Luboshez lens is able to make, in monochromatic light, an

aberration free image of a pointlike object placed in the center of curvature

of its concave surface. This concave surface is metallized and works as the

spherical mirror used in the self imaging TO technique. Radiation from the

fibre is partially reflected by the mirror and coupled back to the fibre itself

to measure P(u). Positioning of the fibre end in the mirror centre is

accomplished observing the image formed by the part of light that crosses L.

This also permits, e.g. in the case of elliptical core fibre, to control the

orientation between the core axis and the scan direction. The aplanatism

condition is satisfied at %=633 nm, because the alignment is performed with an

auxiliary He-Ne laser beam. With component L the mirror NA can be very close

to 1 (value attanaible with a hemispherical concave surface). The observing

optical system can be operated like a common high working distance

microscope .The excitation of the fundamental mode and the measurement of the

radiation coupled back is performed with the same apparatus described in [3].

A fast scan with 1024 points is performed by oscillating L at 60 Hz by a

piezoelectric actuator. The signal from the detector and the sinusoidal signal

(proportional to the mirror displacement) that drives the PZT high voltage

amplifier are sent to the two input channels of a programmable digitizing

oscilloscope, with 1024 points resolution. For an oscillation range of 30 ^im

the average distance between two next points on the P(u) curve is about

0.06 jxm. Thus a high resolution, thanks to the very small scanning step, can

be obtained. In addition a very high SNR is attanaible, owing to the

54
possibility of high speed average. Finally a desktop computer acquires the two

waveforms, provides P(u) and performs the elaboration.

Calculation of MFD . Fig. 2 shows the 80 experimental P(u) central points. The

simplest fitting curve to these data is a parabola, but its parameters and

hence d are heavily affected by the data errors when the number of experimental

points is too small. To increase the number of fitted points it is necessary

to use a larger u-range. But in this case P(u) can depart significantly from a

parabola. However, considering that P(u) is approximately gaussian, we propose

the following fitting function (Fig. 2):

where d

£ is the MFD obtained from the 1/e points of the experimental P(u); A

and B are the parameters to be determined by the fitting procedure. From

Eqs. 1 and 2 one gets d. Fig. 3 shows the calculated d-value versus the number

of fitted points. The first part of the curve, corresponding to few fitted

points, is not reliable. On the other hand for very large number of points the

curve becomes to drift. However, thanks to the chosen fitting function, a

large central range (50-90 points) is left in which the obtained d-value is

relatively constant, allowing a precise determination of the MFD. This could

suggest an effective procedure for MFD evaluation. As a confirmation, a good

agreement with the d-value obtained from a near-field measurement (dashed line)

can also be observed.

(2) P(u) = (A+Bu1

) exp[-(2u/d

e

) 2']

References

1. G. Coppa et al., EFOC/LAN 85, Montreux (CH), 19-21 June 1985.

2. G. Coppa, P. Di Vita, CSELT Rapp. Teen •

»

Vol. 11 (1983), p. 427.

3. M. Calzavara et al Electron. Lett., Vol. 22 (1986), p. 144.

4. M. Herzberger, Modern Geometrical Optics, Interscience , 1958, p. 52.

55
56
Joint Losses as a Function of Offset and Mode Field Diameters for Dispersion-Shifted and

Standard Fibers and Fiber Combinations

V. Shah, L. Curtis, and W. C. Young

Bell Communications Research, 331 Newman Springs Rd., Red Bank, NJ 07701-7020

1. INTRODUCTION

Recently, there has been increased interest in single-mode fibers having zero dispersion at 1550 nm

because the spectral loss of the fiber has its minimum value near that wavelength. These fibers are

called dispersion-shifted fibers, and have more complex index profiles than the standard fibers that have their zero dispersion wavelength at 1300 nm. Contrary to the case of standard fibers, the

fields of such fibers are non-Gaussian in nature. And while the Gaussian approximation has been

shown to give accurate results for the mode field diameter (MFD) of standard fibers for

dispersion-shifted fibers, an alternate definition called "Petermann 2" based on the second-moment

of the far field

^

has recently been shown to give consistent results for commonly used

measurement techniques.^

In this paper, we present the results of the comparison between the measured values of the joint

losses as a function of offset and a theoretical model, based on the Gaussian function of the offset. The comparison is performed on three fibers with different index profiles, two of which are standard matched cladding (MC) and depressed-cladding (DC) fibers, the third being dispersionshifted (DS) fiber with a triangular core with a ring in the cladding region. Both the cases of

identical and non-identical fibers (MFD's different) making up the joint are treated.

MEASUREMENT OF THE MFD:

The MFD of the fibers were measured using the transverse offset method from which the mode

field diameter was determined from the normalized curvature of the transmission coefficient as a

function of offset at the point of zero offset. This definition has been shown to be equivalent to the

second-moment of the far field definition.^ This definition assumes identical fiber on both sides of

the joint. We have extended this definition to include the measurement of the equivalent MFD of

two non-identical fibers in a joint. It is defined to be 2weq = 2/(irW tq ) where, with F\(q) and F2 (g)

representing the far field amplitudes of fiber 1 and 2, respectively,

o (1)

o

and q = — sin 6

57
The equivalent mode field radius (MFR) of non-identical fibers in the joint is then determined by

the normalized curvature as discussed above. The results of these measurements are summarized in Table I. Numbers in parentheses in the table, are values of the equivalent MFR based on the

Gaussian field assumption (i.e., m>£ = V(wf + w£)/2 ). COMPARISON OF THE JOINT LOSS MODEL WITH MEASUREMENT

The joint loss, as a function of transverse offset of fibers with non-Gaussian fields can be predicted

by measuring the far fields and numerically evaluating the inverse Hankel transform'-

4

-'. As an

alternative approach, in this study, a Gaussian function is used to model the joint losses of fibers

(with non-Gaussian fields) as a function of transverse offset. The MFD based on the secondmoment definition (or extension thereof for non-identical fibers) is used to characterize the

Gaussian function used in the model. For standard fibers, loss (dB) due to mis-match = 20

log [(^i+wl)/(2w l w2)], w, and w2 being the mode field radii of the fibers. For DS fibers, an

overlap integral involving the far field amplitudes needs to be evaluated. This overlap integral has

been shown to predict the loss due to mis-match with good accuracy^. In the absence of the far

field data, in this paper, the measured value of the loss due to mis-match is used in the model.

Figures 1 and 2 show theoretical and measured joint loss vs offset at 1300 and 1550 nm for the MC

and DS fibers, respectively. While the agreement between the measured and theoretical values is

excellent for the standard fiber, for the DS fiber, there is good agreement for losses up to about

2 dB. For losses greater than 2 dB, the model predicts higher losses compared to the measured

values.

The loss vs offset plot for MODS combination is shown in Fig. 3. It is evident that the model,

using modified "Petermann 2" MFR (wtq) predicts joint losses with good accuracy at both

wavelengths. Likewise, the same performance was observed for DC/DS and DC/MC combination.

Figures 4 and 5 show the loss vs offset plots at 1300 and 1550 nm, respectively, for MC fiber

(curve A), DS fiber (curve B) and MC/DS combination (curve C). These three curves are measurement results. Curve D in both figures represents the calculation for MC/DS joint based on

the theory developed in reference [6]. The loss in dB for the offset d is given, in this case, by^

Lots=2o,°g [i!^f]

+ 4- 343 (4r <2)

where w l and w2 are measured MFR of individual fibers and w?q

is taken from the numbers in

parenthesis in Table I. It is clear that this model overestimates the losses. The difference between

the measured and calculated values, curves C and D for MC/DS joint, is smallest at zero offset and

increases as the offset increases. The same performance holds for DC/DS combination.

58
CONCLUSION

For both the cases of identical fibers (standard and dispersion-shifted) and DS/Standard fiber

combination, a Gaussian joint loss model, characterized by the MFR based on the second-moment

of far field (or its extension for non-identical fibers), was found to predict joint loss (at 1300 and

1550 nm), as a function of offset, within the range of practical interest with good accuracy. The

difference between the measured and the theoretical loss value was less than .05 dB at 2 dB.

The authors would like to thank Corning Glass Works for providing the dispersion-shifted fiber

used in this study.

REFERENCES

[1] W. T. Anderson, IEEE J. of Lightwave Tech., Vol. LT-2, No. 2, p. 191, 1984.

[2] K. Petermann, Electron Letts., Vol. 19, No. 18, p. 712, 1983.

[3] W. T. Anderson, et al., to be submitted for publication.

[4] C. Pask, Electron Letts., Vol. 20, No. 3, p. 144, 1984.

[5] W. T. Anderson, et al., submitted for publication.

[6] D. Marcuse, BSTJ, Vol. 56, No. 5, p. 703, 1977.

TABLE I

Wavelength Fibers

MC DC DS DS/MC DS/DC DC/MC

1300 nm 4.95 4.50 3.33 4.45

(4.22)

4.3

(3.96)

4.72

(4.73)

1550 nm 5.70 5 AS 4.26 5.31

(5.03)

4.95

(4.72)

5.37

(5.43)

Offset (nm) Offset (\im)

Fig. 1 Joint loss vs offset plot for MC fiber. Fig 2 Joint loss vs offset plot for DS fiber.

Solid lines are theoretical calculations. Solid lines are theoretical calculations.

• and A represent measured values. • and A represent measured values.

59
Offset (urn)

Fig. 4 Comparison of measured joint loss values with the theoretical model based on

Eq. (2) at 1300 nm for MC/DS combination. Curves A,B and C are measurement

results while D is theoretical curve.

60
A NOVEL SINGLE-MODE FIBER SPLICE

MEASUREMENT USING LOCAL

ALIGNMENT AND LOSS

DETECTION

Ian A . White and Paul D.

AT&T Bell Laborator

2000 Northeast Expre

Norcross, Georgia

Lud i ng ton t

i es ssway

30071

(404) 447-2696

A mandrel-wrap local detector (MWLD) is described that is used to

align single-mode (S-M) fiber splices to within 0.2 vim and to

accurately measure splice loss. A simple theoretical result for

splice loss prediction from scattered power measurements is given

and generally predicts splice loss within 10$ when compared to

far-end transmission measurements. Field results for some 10,000

splices, using this method, have yielded average losses of < 0.05

dB.

INTRODUCTION

Local detection of power lost in S-M fiber splices can be used to

assist in alignment of fiber cores during splicing and to determine splice loss. Many authors have reported the use of such

local detection schemes. 1_lt The estimation of splice loss by

these schemes is generally based on calibration of the splice

loss versus detected power, and can be in error due to variations

in the parameters of the receiving fiber. In this paper, we present a local detection system for the alignment of S-M splices,

and a new theoretical calculation (using this locally detected

scattered power) to accurately predict splice loss. This

technique has been incorporated in the AT&T Technologies SingleMode Bonded Splicing Apparatus. 5 More than 10,000 such splices

have been successfully completed in the field with average losses

of less than 0.05 dB.

LOCAL DETECTION SYSTEM

In a S-M fiber splice, the power lost due to the splice is

coupled into the radiation modes of the receiving fiber. By

selectively detecting only this scattered power, without attenuating the fundamental mode, the fiber cores can be accurately

aligned using the criterion of a minimum detected signal (null).

In Figure 1, we illustrate the detection scheme using the mandrel

wrap local detector (MWLD). A mandrel (radius 1.6 cm) is chosen

to tap selectively the scattered power (from the splice) out of

tNow with Kodak Inc., Rochester.

61
the fiber, in the vicinity of the detector, without inducing anysignificant loss of the core power. A large area Ge detector (5 mm dia.) located tangential to the mandrel is used to detect the

scat ter ed power . Highly accurate stepping motors (0.1 ym step size) are used to

align the fiber cores in both the horizontal and vertical

directions wi th feedback provided by the MWLD. These accurate

stepping motors are necessary for the loss evaluation discussed

below. Figure 2 shows a typical scan profile in the horizontal

direction for such a system superimposed on the refractive index

profile of the receiving fiber. The sharp null, which is located

at the center of the symmetric scan profile, corresponds directly

to optimum splice alignment. (Note that the shape of the scan

profile is related to the structure of the receiving fiber.)

THEORY OF LOCAL SPLICE LOSS MEASUREMENT

From the microbending theory of S-M fiber splicing, 6 more than

90% of the power lost from a splice with small amounts of offset

and tilt, is coupled into the LP-|i mode. By measuring the scattered power level in the null, P-|, and the power level, ?2t at

some known offset from optimal alignment, r this power is in the same radiation mode

tions it is easily shown that the splice

aligned position (null) is

0 , and knowing that

for both offset posiloss at the optimally

Loss (dB) = 4.3^

1 -P

( 1 )

The effect of the variability of detected power levels due to

fiber design variations and differences in the input power to the

splice are eliminated since only the ratio of the detected power

levels P is used in the loss calculation.

P = (2)

The only constants in Eq.

nominal mode field radius

the loss calculation is a

( 1 ) are the

(w Q) of the

function of

of w Q , which is typically well within

known offset r 0 , and the

LPq1 mode. The error in

the variations in the value

8? of nominal.

62
SPLICE LOSS RESULTS

In Figure 3» the calculation of the splice loss of an optimally

aligned splice using equation 1, is shown as a function of the

choice of induced offset, r 0 , from Optimal alignment. For compar

ison, the loss of this splice using a conventional far end

transmission measurement is also shown. For r 0

> 2 vim the perturbation analysis used to derive equation 1, begins to fail,

resulting in erroneous predictions. Figure 4 shows a comparison

of local loss predictions to far end measurements for the practical range of single mode fiber splices of 0 to 0.5 dB, illustrating that the local loss predictions are generally within 10% of

the far end measured losses.

APPLICATIONS AND CONCLUSIONS

This alignment technique and measurement procedure has been used

in the AT&T Technologies Single-Mode Bonded Splicing Apparatus,

for over 10,000 field splices with average losses of less than

0.05 dB. This technique is independent of the individual fiber

parameters and accurately predicts loss.

REFERENCES

1. C. M. Miller, Proc. OFC 1981, THAA2.

2. K. Kato, T. Tanifuji, M. Tokada and N. Uchida, Electronic

Letters J_8 , 972 ( 1 982 ).

3. T. Haibara, M. Matsumoto, T. Tanifuji and M. Tokuda, Optics

Letters 8 , 235 ( 1 983 ) .

4. C. M. De Blok and P. Matthijsse, Electronic Letters 2_0, 109

(1984).

5. M. R. Reynolds and P. F. Gagen, Proc. I.C.C., Amsterdam 1984.

6. I. A. White and J. F. Kuhl, Proc. I.O.O.C., Tokyo 1983-

63
Refractive I n d e

:

U3 in m

- o

1*13 O

C 0 1 l. U Q.

1/1 X

u o U Tl 3 C O

o. B

"D > V —

C 13

•-i C

L. <0

Orco n m m

}nd}.no pszi [Eujjom

orO

! r i

64
FUSION SPLICING OF DISSIMILAR FIBERS- A COMPARISON OF MODE FIELD DIAMETER

AND CROSS-CORRELATION LOSS PREDICTIONS WITH EXPERIMENTAL RESULTS

by W. T. Anderson, A. J. Johnson, J. P. Kilmer, and E. A. Thomas

Bell Communications Research

1. Introduction

In this paper, we report results from splicing single-mode fibers of different designs and attempt to

interpret the results in terms of the degree of optical and physical compatibility between the fibers. When fusion splicing two single-mode fibers, a number of factors can contribute to the loss of the joint.

While many of these factors relate to the fusion splicing machine and its operation, one factor over which the operator has no control is the degree of optical and physical compatibility of the two fibers. A wide range of designs are commercially available (matched and depressed cladding, conventional and

dispersion-shifted) and two processes, each with several different implementations, are used to make the

fibers (inside and outside deposition). Both design and process differences can affect the compatibility of two fibers with respect to fusion splicing.

Losses for butt-joint splices have been successfully predicted by the mode field diameter mismatch

between the two fibers [l]. However, this single parameter, which defines the width of the fields, is adequate only if the field shapes of the two fibers are nearly identical. This limitation applies to both

the gaussian mode field diameter definition [2] and to the recently proposed second-moment definition

[3,4]. A more complex model, which is based upon the cross-correlation of the fields of the two fibers,

has been used to predict butt-joint losses for fibers of dissimilar designs [5]. In this paper, loss measurements for fusion splices between three combinations of substantially dissimilar fibers are compared to three splice loss predictions: one based upon mode field diameter measurements with a gaussian definition, another based upon mode field diameter measurements with a second moment

definition, and a third based upon field cross-correlation.

2. The Experiment

The electric arc fusion splicing machine used was microprocessor-controlled, and the arc settings were

determined by trial and error by an experienced operator. Alignment of the two fiber ends was done

visually using a 120x microscope. The alignment stages have minimum resolution of O.l^m. Power

monitoring was not used. An arc time of 2.0 seconds, which allows for some self-alignment due to

surface tension, was used for all splices. When used with fibers with nearly identical mode field radii,

this technique produced splices below .02dB.

The splices involved four commercially available fibers. Three of them were conventional step-index 1

single-mode fibers: a depressed-cladding inside process fiber ("DI"), a matched-cladding inside process

fiber ("MI"), and a matched-cladding outside process fiber ("MO"). The fourth fiber was a commercially-available dispersion-shifted fiber [6] produced by the outside process ("SO"). Three

combinations of these four fibers were studied: SO to MO, DI to MI, and SO to MI.

65
Two OTDRs were used to make initial loss measurements at both 1300 and at 1550nm. Because of the

differences in fiber designs and the corresponding differences in backscatter capture coefficients, a

difference in the OTDR readings taken in different directions was expected, but the average of the two

OTDR readings taken in opposite directions gives an accurate measure of the splice loss. A series of

splices were made and measured to determine the repeatability of the splices, and variations of less than

0.03dB were typical. After several splices, a joint whose loss was on the low side of this distribution was

selected, and the sample was saved for a transmission measurement of the loss using the cutback

technique. Enough slack fiber was saved so that the cutback measurement could be repeated twice in each direction, and the overall accuracy of the transmission measurement is estimated to be .02dB. The

splice loss measurements for the three fiber combinations are shown in the table below.

Measured Splice Losses (+) or Gains (-) in dB

Fiber Wavelength OTDR Measurement Transmission Measurement

Combination turn) Fwd Rev Avg Fwd Rev Avg

SO - MO 1300 2.40 -1.47 .46 .413 .517 .46 1550 1.73 -1.20 .26 .232 .263 .25

DI - MI 1300 0.93 -0.57 .18 .204 .199 .20 1550 0.96 -0.62 .17 .201 .202 .20

SO - MI 1300 3.53 -1.68 .92 .973 .908 .94 1550 2.65 -1.59 .53 .545 .464 .50

The averaged OTDR and transmission measurements agree well. The large directionality of the OTDR

measurements was expected. However, the significant directionality observed for the transmission

measurements of the first and third combination is surprising.

3. Splice Loss Prediction from Mode Field Diameter and Cross-Correlation

The mode field diameters of the four fibers were measured at the two wavelengths using the far field scan method [2]. Both the gaussian and the second moment definitions were used. The estimated splice

loss is found from the two mode field diameters MFD^ and MFD2 as [l]

20 log

MFD? + MFDl

2 MFDi MFD2 dB

The splice loss can also be predicted from the more complex cross-correlation, or overlap, of the fiber

fields. Because of the Hankel transform relationship between the fiber near fields /(r) and the fiber far

fields F(q), the near-field cross-correlation and the far-field cross-correlation are identical [7], so the

splice loss is predicted by

66
loss = lOlog

12

Jfi(r)f 2(r)r dr

10 log

lFx {q)F2{ q )q dq

o dB

Jff(r)rdr //2

2 (r)rdr {Ff(q)qdq J F§(q)q dq

where (7 = -^-sin0 and the maximum scan angle 9 m is chosen so that qm = -^-sin0 m is large enough to X X accurately calculate the above integrals. The far field measurements were conducted over a range of ±25 0

, but the 60dB dynamic range of the system limited the range of useful data for the step index

fibers to about ±20° . The far fields of each of the four fibers were measured twice, and the average of those two measurements are summarized below. The standard deviations were typically .04//m.

Measured Mode Field Diameters (in nm)

Fiber Gaussian MFD

1300nm 1550nm

Moment MFD

1300nm 1550nm

SO 6.82 8.86 6.54 8.18 MO 9.57 10.95 9.51 10.83

MI 11.07 12.21 10.70 11.83

DI 9.09 10.29 8.93 10.16

The MFD and cross-correlation predictions are compared to the measured splice losses below.

Splice Loss Predictions and Measurements fin dB)

Fiber Wavelength Predicted Splice Loss Measure d Splice Loss Combination (im Gaussian Moment Cross-cor OTDR Trans

SO - MO 1300 0.49 0.59 0.43 0.46 0.46

1550 0.19 0.33 0.20 0.26 0.25

DI - MI 1300 0.17 0.14 0.16 0.18 0.20

1550 0.13 0.10 0.10 0.17 0.20

SO - MI 1300 0.98 1.01 0.90 0.92 0.94

1550 6.44 0.57 0.45 0.53 0.50

The differences between the three predictions and the two measured values are shown in the following

table. Positive values indicate that the prediction is lower than the measured loss while negative values

indicate the opposite.

Differences Between Loss Predictions and Measured Losses (in dB)

Fibers Wavelength

fxm

Gaussian MFD

OTDR Trans

Moment MFD

OTDR Trans

Cross-correlation

OTDR Trans

SO - MO 1300 -.03 -.03 -.13 -.13 .03 .03 1550 .07 .06 -.07 -.08 .06 .05

DI - MI 1300 .01 .03 .04 .06 .02 .04 1550 .04 .07 .07 .10 .07 .10

SO - MI 1300 -.06 -.04 -.09 -.07 .02 .04 1550 .09 .06 -.04 -.07 .08 .05

The cross-correlation prediction is systematically lower than the measured loss by 0.02 to O.lOdB. Since

67
these three fiber combinations are expected to be incompatible physically as well as optically, the

predicted splice loss, which includes the effect of only optical incompatibility, should be somewhat

smaller than the measured splice loss. Also, extrinsic factors, such as misalignment of the two fibers,

will tend to increase the actual loss above the prediction. Considering that nearly identical fibers produced as much as 0.02dB loss using the same splicing machine and procedure, the cross-correlation

prediction is reasonably consistent with the measured values. On the other hand, both of the mode field diameter predictions are neither consistently above nor below the measured values. Since a predicted

loss which is higher than the loss actually observed (which would be negative in the above table) cannot

be explained by extrinsic factors or physical incompatibilities, which can only increase the loss, these estimates are inconsistent with the measured values.

The predictions of the cross-correlation model are very consistent for the three splices. Since the first two combinations both involve fibers made using the same process and using similar dopants, these combinations were expected to be more physically compatible than the third combination, which

involves two fibers from different processes and contain considerably different materials. If the different processes and materials significantly affected the fusion process, then the predictions, which include only

optical compatibility considerations, would be worse for this combination than for the other two

combinations. This expectation was not realized, and optical compatibility alone is sufficient to explain

most of the observed loss. Also, the cross-correlation prediction is closer to the measured values at 1300nm than at 1550nm (average differences of .03dB verses .07dB). The reasons for this systematic

difference are not known.

4. Conclusions

The mode field diameter, which is a simple one parameter description of the field distribution, is not expected to accurately predict the loss of splices between fibers whose field shapes are substantially

different. A more complex model seems to be necessary to predict the loss accurately. Additional work

is required to verify the accuracy of this model with other fiber designs and to more fully understand the

directionality which some splices exhibit. Also, a comparison of butt-joint splices with fusion splices

could confirm the conclusion in this study that physical incompatibility contributes relatively little to

the loss of fusion splices.

The authors wish to acknowledge the contribution of Corning Glass Works, who provided the sample of

dispersion-shifted fiber used in this study.

REFERENCES

[1] D. Marcuse, BSTJ, Vol. 56, No. 5, p. 703, 1977.

[2] W. T. Anderson and D. L. Philen, IEEE JLT, Vol. 1, No. 1, p. 20, 1983.

[3] K. Petermann, Electronics Letters, Vol. 19, No. 18, p. 712, 1983.

[4] C. Pask, Electronics Letters, Vol. 20, p. 144, 1984.

[5] J. C. Auge and L. B. Jeunhomme, Technical Digest, SPIE Tech Symp, August, 1985.

[6] T. D. Croft, J. E. Ritter, and V. A. Bhagavatula, Technical Digest, OFC 85, Paper WD2, 1985.

[7] W. T. Anderson, IEEE JLT, Vol. LT-2, No. 2, p. 191, 1984.

68
STRESS-INDUCED INCREASES IN INTERFERENCE EFFECTS AT FIBER JOINTS AND A

METHOD TO ELIMINATE INTERFERENCE EFFECTS IN OPTICAL FIBER

MEASUREMENTS

W. C. Young, V. Shah, and L. Curtis

Bell Communications Research, 331 Newman Springs Rd., Red Bank, NJ 07701-7020

INTRODUCTION

A common practice used in characterizing single-mode fiber systems, such as laser power

measurements, bit-error-ratio measurements, and spliced fiber/cable loss measurements, usually

employs the use of a "photon-bucket" (PB) type receiver jumper cable to access the fiber/cable ends.

Typically, these PB jumpers consist of a length of multimode fiber with a sufficiently large core to

collect all of the transmitted power from a single-mode fiber without regard to precision alignments.

However, if other characteristics of these fiber joints are not controlled, a significant uncertainty in

the measurements can be encountered. Our measurements show that for repeated measurements

this uncertainty can result in loss variations as high as 1.14 dB. In this paper we describe the source of these loss variations and discuss techniques that can be used to minimize resulting measurement

uncertainties.

DISCUSSION

Practical consideration in measuring the optical power transmitted by a single-mode fiber dictates

the use of a large-core fiber (multimode fiber) as a "photon-bucket" receiver. Using this MM fiber

avoids the uncertainty, due to the required submicron alignment accuracies, in the loss of SM/SM

fiber connectors. Furthermore, to increase their engagement/separation endurance level these

SM/MM connections are designed to establish fiber endface separation. As a result of this

separation and reflections from the fiber end faces, a Fabry-Parot etalon is formed at the joint.

When excited with a coherent source, multiple-beam interference occurs, as shown in Figure 1. The resulting interference effects cause large variations in transmitted power, as a function of fiberto-fiber separation, as long as the separation is smaller than the coherence length of the source. With LEDs, the effect is significantly diminished with a separation of about 10 to IS micrometers,

whereas for a highly coherent source such as a laser, it persists for very large separations on the

orders of 100 p.m, and beyond. Since a large percentage of measurements are made with laser

sources, these effects can cause significant measurement uncertainties. Experimental measurements

of these interference effects for SM/SM joints, using both a laser and a LED, are shown in Figure 2. The decrease in the mean transmitted power, with increasing separation, is due to the aperturing

effect of the SM receiving fiber. In Figure 3 similar measurements are shown, but in this case we

have used a SM/MM fiber joint. The receiving MM fiber was a 50/125 GI fiber. As can be seen in

these figures, the variation in transmitted power, with laser sources, can be significant even for

substantial separations. For the experimental results shown in Figures 2 and 3, the fiber endfaces

69
were polished, flat and perpendicular to the fiber axis within .2 degrees, using a one micron

aluminum-oxide coated sheet of polyester. The polishing was performed dry with no lubrication.

As previously reported, the maximum loss variation for air-gap joints having flat and parallel faces can be calculated from:

variation = -10 log (l-4rt) (l)

where n is the refractive index of the fiber.

For a refractive index of 1.46, typical for a fiber, theory predicts a maximum loss variation of 0.6S

dB. However, experimentally, maximum variations of 0.90 to 0.95 dB were measured (Figures 2

and 3). To observe these higher values, the refractive index of the polished endfaces would have to be about 1.57. The possibility of the presence of a thin film of aluminum-oxide, which has a

refractive index of about 1.6, after polishing was investigated with an SEM. However, no traces of

such a film were found. Another cause which is more likely the reason for the high variation is that due to stresses induced during polishing the density of the glass surface increases, thus causing an

increase in the refractive index. Previous work has shown that when silica glass was polished with

red iron-oxide the index was increased from 1.46 to about 1.6^. Investigating the effect that

different polishing techniques have on the maximum variation in loss (maximum change in the index

of the surface), we were able to achieve a maximum variation of 1.14 dB by changing the rate and

amount of polishing, while still using one micron aluminum-oxide grit. This loss would require a

surface index of about 1.63. In contrast, when we used cleaved perpendicular faces, the maximum

variation was 0.6 dB, which is expected for the refractive index of fused silica, namely 1.46.

Regardless of what effect the surface condition has on the index, a method to minimize the

variations due to the interference effect is needed. We have achieved this minimization by

introducing a face angle on the MM receiving connector. Measurements of transmitted-powerversus-separation for SM/MM joints, for MM fibers have various face angles are shown in Figure 4. The MM fiber was a 50/125 GI fiber. Based on these results, an appropriate face angle can be

introduced, on the connector end, which significantly minimizes the interference effects in air-gap

connectors. An alternate solution would be to use an index matching material in the joint but

generally this approach has not been accepted as a practical solution due to problems such as

contamination, handling, etc. 1 J

70
In an experiment, we have used SO/125 GI fiber receiver jumpers, with a 5 degree face angle and a gap of about 25 microns, to measure end-to-end loss of 5 concatenated sections of SM fiber (about

35.7 km). The coupling loss of the SM/MM connectors ranged from 0.62 to 0.65 dB for repetitive

measurements, in contrast to the range of 0 to 0.95 for parallel polished endfaces. For repeated

end-to-end loss measurements, made in both directions through the fibers, loss measurements at 1300 nm ranged from 15.86 to 15.91 dB.

CONCLUSION

We have shown that interference effects in air-gap connectors can be affected by stresses induced in

the fiber endface during polishing, and that by employing a face angle of 5° on a MM receiving test

fiber, the maximum variation due to these interference effects can be reduced from .92 dB (0° end

face angle) to less than .05 dB. Finally, since a large reduction in the interference effect can easily

be achieved by using the endface angle on MM fiber, the accuracy of optical fiber measurements can be substantially improved.

REFERENCES

[1] W. C. Young et al., "A Transfer Molded Connector Biconic Connector with Insertion Loss

Below 0.3 dB Without Index Match", Technical Digest, Sixth ECOC, U. York, 1980.

[2] R. E. Wagner and C. R. Sandahl, "Interference Effects in Optical Fiber Connections",

Applied Optics, Vol. 21, No. 8, pp. 1381, 15 April 1982.

[3] L. Rayleigh, "The Surface Layer of Polished Silica and Glass with Further Studies of Optical

Contact", Proc. Roy. Soc, A 160, pp. 507, (1937).

[4] I. Sankawa et al., "Methods for Reducing the Fresnel Reflection in an Optical-Fiber

Connector with Index Matching Material", Elect, and Comm. in Japan, Part 1, Vol. 69, No.

1, pp. 94, (1986).

Fig. 1 Multiple-beam interference at a fiber joint

71
Fiber Separation (MICRONS)

Fig. 2 Measured transmission fluctuation

vs. fiber end face separation for a SM/SM joint.

Fiber Separation (MICRONS)

Fig. 3 Measured transmission fluctuation vs fiber end face separation for a SM/MM joint.

iodu iwwwww^^ mm

.04dB

0 25 80 75 100

Fiber Separation (MICRONS)

Fig. 4 Measured transmission fluctuation vs fiber end

face separation for a SM/MM joint with a 1300 nm

laser. The MM fiber end face in the joint is polished at different angles.

72
PRACTICAL ONE-WAY-LOSS RANGE FOR OTDRs

M C Brain and E A Cottrell

British Telecom Research Laboratories

Martlesham Heath, Ipswich IP5 7RE , UK.

Introduction

The method of detecting remote breaks in fibre transmission systems, bymeasuring backscattered optical power as a function of distance along the

fibre [1] , has made the Optical Time Domain Reflectometer (OTDR) an indispensible test instrument. In this paper, we show that the range of fibre

loss over which breaks may be accurately located is less than the

conventionally defined one-way-loss range. Our discussion is illustrated with

measurements at 1.5 /j,m over 100 km of commercial single mode fibre.

Theory

An OTDR is shown schematically in Fig 1, and we develop the theory as follows.

An optical pulse of peak power Ps from the laser is launched into the fibre

via a directional coupler, and after a delay r a backscattered signal Px

is

detected from a mean distance x into the fibre. The difference between P s and

Px

is determined by the total loss C in the coupling optics, the fibre loss

2Lx incurred over the distance x on the outward and the return paths, and the

reflection coefficient R for the signal backscattered in the vicinity of x within the capture angle of the fibre core [2, 3]. Thus,

2Lx = P s

- (R+C) - Px

dB (1)

The maximum range over which measurements can be made is reached when Px approaches the noise equivalent power (NEP) P<j of the detector, although this can be improved by a factor 51ogioN ( = SNIR/q [3]) by averaging N

consecutive measurements. We thus arrive at a definition for the one -way- loss

range L of an OTDR [3]:

L = [P s

- (R + C) -Pd + SNIR/q ]/2 (2)

This definition is convenient for comparing the performance of different OTDR

instruments, but for practical measurements it must be noted that a limiting

signal-to-noise ratio (SNR) of unity is implied by defining L in terms of the

NEP of the detector. Close to this limit, where there is a high probability

73
of the signal-plus-noise being near zero, it would be impossible to locate

accurately the position of a break giving little or no Fresnel reflection,

since the logarithmic signal scale of a conventional backscatter trace would

yield large negative -going spikes, which would mask the abrupt fall in signal

expected at the break. The maximum actual fibre loss must therefore be less

than L, by say M dB. The question arises, how large must M be?

If in (1) is used to represent the new one-way-loss range, then

M = L - Lx = [Px

- Pd + SNIR/q]/2 = 51og10 (Ix/a ) (3)

where I x

is the postdetection photocurrent due to the backscattered signal

from x, and a is the postdetection RMS equivalent input noise current of the

receiver. Thus Ix = We now estimate the amplitude AL of noise

fluctuations on the backscatter trace by considering an envelope of

+2ct about the mean postdetection photocurrent; for Gaussian white noise,

this envelope includes over 95% of the noise distribution. By a derivation

similar to (3) , we find:

A L = 51og10 [(Ix ± 2a )/Ix ] = 51og10 (l + 2.10" M/ 5

) (4)

This is plotted in Fig 2 as a function of M, suggesting that a value of M

greater than 1.5 dB is essential if large negative -going spikes are to be

avoided, but a margin of 3 dB would reduce the -la deviation to -1.5 dB.

Measurements

We now compare these predictions with practical measurements using a 1.5 jj.m OTDR [4] . A semiconductor laser was used to launch 5 jj,s pulses of peak

power 7.6 mW at 1.535 /j,m wavelength into commercial monomode fibre using a

3 dB single-mode fibre fused taper coupler. The fibre consisted of 19

lengths, fusion spliced to give a total length of 101.2 km with loss 22.8 dB.

The receiver comprised a commercial InGaAs PIN photodiode (75 /xm diameter,

2.8 nA leakage, 1.08 A/W responsivity) , followed by a transimpedance

preamplifier with a silicon JFET input stage and a 10 Mohm resistor in the

feedback loop. Single channel averaging improved the SNR to yield an expected

one-way-loss range L = 16.3 + 2.51ogioN dB, obtained from (1).

The implications of Fig 2 are demonstrated in Fig 3, which shows the

Via levels plotted about the measured backscatter traces for different

74
values of N. A one -way- loss range equal to the known fibre loss requires that

400 samples be averaged at each measurement point. The trace in Fig 3a was obtained with 1500 samples/point, suggesting a margin M of 1.4 dB, yet the

location of the fibre end is obscured by the noise. When N is increased to

7500 samples/point (Fig 3b), the margin rises to 3.2 dB and the fibre end

becomes distinct. Further increasing N to 75000 samples/point (Fig 3c)

allows the fibre end to be characterised to an accuracy of +0.3 dB

(+2ct levels), corresponding to a margin M of 5.7 dB. In all cases the

la envelope is a convincing fit to the measured trace. Furthermore, 7500

samples at a repetition rate of 900 Hz gives a measurement time of 8.3

seconds/point , which is acceptable for practical situations.

Conclusion

We have demonstrated a break location capability over 100 km of commercial

single mode fibre, using an OTDR operating at 1.535 ^tm wavelength, assembled

from commercial components. We have used these measurements to demonstrate

that for accurate fibre end location in the absence of Fresnel reflection, the

actual fibre loss should be ~3 dB less than the one -way- loss range defined in

terms of unity SNR. A margin of 5 dB allows the entire fibre to be

characterised to better than +0.5 dB.

Acknowledgements : the authors thank the Director of Technology Applications,

British Telecom, for permission to present this paper.

References

[1] C Y Boisrobert, Proc Topical Meeting on Optical Fiber Transmission,

January 7 - 9 1975, Williamsburg Va, USA.

[2] E Brinkmeyer, Electronics Letters, Vol 16 No 9, pp 329-330, 1980.

[3] P Healey, J Lightwave Technology, Vol LT-3 No 4 , pp 876-878, 1985.

[4] E A Cottrell, M C Brain, Electronics Letters, Vol 22 No 8, pp 443-445,

1986.

75
101.2 km

test fibre

LASER

InGaAs PIN RECEIVER

\

index-matched

fibre-ends

SAMPLER A/D

DELAY

GENERATOR

H P

SCOPE XY

PLOTTER

Fig 1. Schematic of OTDR equipment.

2 3 4 5

Margin M, dB

Fig 2. Deviation from mean backscatter level as a function of margin M between one-way-loss

range L and actual fibre loss.

CD

•om

inO

-15r

° -20 - o

I 4)CO

-25 -

-30 -

700 800 900 1000

Loop delay time, ps

fibre end

700 800 900 1000

Loop delay time, ps

m

T3m

(0 _oo

i

I <DcO

fibre end

-25

800 900 1000

Loop delay time, ps

Fig 3. Plots of backscatter signal from the remote

fibre end. The margin between the estimated one-way- loss range of the OTDR

and the actual loss of the test fibre is: (a) 1.4 dB (unacceptable)

(b) 3.2 dB (Just acceptable for fault location)

(c) 5.7 dB (accuracy +0 3 dB)

76
A NEW TECHNIQUE FOR THE RELATIVE

MEASUREMENT OF SCATTER LEVELS

IN SINGLE MODE FIBRES

M.E.Fermann, S.B.Poole, D.N.Payne, F.Martinez

Optical Fibre Group, Department of Electronics and Engineering, Southampton

University, Southampton S09 5NH

ABSTRACT

A new sensitive technique to separate scatter loss contributions from total

fibre loss in single-mode fibres based on Optical Time Domain Reflectometry

OTDR is described. Results for telecommunications-grade fibres with

different GeC^ concentration in the fibre core are presented.

INTRODUCTION

The ability to measure relative scatter levels in single-mode fibres is

essential for minimising fibre losses in telecommunications fibres operating

in the 2nd and 3rd windows for optical communications. In this way

fabrication processes can be optimised in response to small improvements in

measured scatter levels. Surprisingly, there is no simple way in which this

can be done with sufficient accuracy to suit current very-low fibre scatter

losses.

We, therefore, describe a new technique which allows the separation of the

Rayleigh scatter loss from the total loss in a single mode fibre. In this

method, a reference fibre, the scatter loss of which may be determined by

calorimetric techniques, is spliced to the fibre under investigation. The

two fibres are then probed using OTDR. The backscattered light powers from

each side of the splice are compared by using the test and reference fibres

in turn as the launch fibre and from this the relative scatter losses of the

fibres may be obtained. The technique is capable of determining very small

changes in scatter from fibre to fibre and is, therefore, ideally suited to

process control and optimisation.

77
THEORY

It may be shown from ref.1 that, in the case of the single-mode step-index

fibres operating around second-mode cutoff, the ratio of the scatter loss

spot sizes here are defined as the (1/e ) power diameter of the Gaussian

approximation to the power distribution of the fundamental mode in the

fibre . The ratio of the scatter losses between a test and a reference

fibre may, therefore, be readily obtained from measurements of the relative

backscattered power, fibre spot sizes and core refractive indices.

MEASUREMENTS

The measurements were carried out at 1 . 3 u m on a number of single-mode

communication grade step index fibres, the parameters of which are

summarised in Table 1 . These fibres all had a GeO^-doped core and a

matched silica cladding, with the exception of fibre 5 which had a slightly

depressed cladding. The fibres were fabricated using VAD and MCVD

techniques. As a standard we used fibre 1 , a low-loss fibre that exhibited

a total attenuation of 0.39± .02 dB/km at 1.3 um. The backscatter traces

were obtained using commercial equipment operated with a resolution of 40

and 75 metres. The spot sizes were measured using a transverse offset

technique.

RESULTS AND DISCUSSION

The ratio a j-/ a s is plotted in Fig. 2 as a function of numerical aperture NA.

The accuracy obtainable for these ratios is limited by the accuracy of the

spot size measurements. The resulting errors in the scatter loss ratios

coefficients a . and « s of a test and a standard fibre is given by

a , / \ 2 / nJ_\ 3 / R- \ U

where R^= P^/P^and R

2

= P

2i/p22 are the ratios of ^he backscattered powers

at the splice as defined in Figure 1 ; n

fc and n

g are the core refractive

indices, and w^.,w

g are the spot sizes of fibre t and s respectively. The

78
were about 5% and are indicated in the Fig. 2 with error bars. Assuming a

scatter loss of silica of 0.3 dB/km at 1.3um, differences of less than 0.015

dB/km in scatter loss may be detected.

4

As expected, the Rayleigh scatter loss increases proportional to NA and

this goes some way to explain the commonly-observed higher attenuation of

fibres with high NA. Surprisingly, the manufacturing process does not have

a measurable effect on scatter loss. The measured variations in total

fibre losses are greater than would be expected solely due to changes in NA

and are due to OH~ peaks and process imperfections.

The measurement technique described here may be applied to any type of

single mode fibre, provided an expression for the relative backscatter

capture fractions of the fibres is computed. Thus triangular core,

depressed-clad or even W-fibres could be measured.

CONCLUSIONS

We have described a new sensitive technique for the measurement of relative

scatter losses in single-mode optical fibres. As an indication of the use

of the method, results showing the measured increase in scatter loss with

increase in NA have been presented.

ACKNOWLEDGEMENTS

The authors would like to thank Pirelli General pic and York VSOP for

supplying fibres. We would also like to thank York Technology Ltd. for

valuable assistance with the spot size measurements. The work was carried

out under the JOERS programme. A readership (DNP) and Research Fellowship

(SBP) were provided by Pirelli General pic.

79
REFERENCES

1 . A.H.Hartog and M. P. Gold. "On the theory of backscattering in

single-mode optical fibres." Journal of Lightwave Technology,

Vol.LT-2 (2), 76-82, 1984.

2. D.Marcuse. "Gaussian appoximation of the fundamental modes in

graded index fibres", J.Opt.Soc. , Vol.68, pp. 103-1 09, 1978.

3. York Technology FQn 1000.

4. H.Murata, S. Inao, Y.Matsuda, and T.Kuroha, "Optimum design for

optical fibre used in optical cable system", Proceedings, 4th

European Conference on Optical Communications, Genoa,

pp. 242-248, 1978.

Fibre Spot- Cut- Total at/ NA ManufacNumber size off loss a s turing

pm nm dB/km process

1 11.0 1250 0.39 1 0.10 VAD

2 9.8 1250 0.54 1.08 0.12 VAD

3 9.5 1250 0.54 1.10 0.12 VAD

4 8.5 1200 0.6 1 .09 0.13 MCVD

5 6.3 1250 0.7 1 .23 0.17 MCVD

6 7.6 1200 0.5 1 .15 0.15 MCVD

Table 1 . Fibre parameters.

Launch

end _

Fig.l Definition of

Parameters

.

80
MODE FIELD DIAMETER EFFECTS ON OTDR SPLICE MEASUREMENTS

Felix Kapron, Carrie Kozikowski, and Ricky Crotts

ITT Electro-Optical Products Division

7635 Plantation Road

Roanoke, VA 24019-0065

Abstract

It is shown that OTDR measurements will give erroneous

results on splice or connector losses that are due primarily to a difference in the mode field diameters of the joined fibers. A

new simple formula gives an improved estimate of loss.

Splice Real Loss and OTDR

Within a splice or connector, let 2w^ and 2w2 be the mode

diameters of the input and output fibers, respectively. Define a relative mismatch R equal to the ratio of the mode diameter

difference divided by the mode diameter average:

R = 2(W! - w 2 )/(wi + w 2 ) (1)

The resulting power transmitt ance [1] can be rewritten as a loss

dB(R) = 20 log 10 (o.5 [(w]/w 2 ) + (w 2 /wi)])

V J (2)

= 4.34 3 R2 for |r| < 0.8

By convention here, an OTDR signal drop is considered

pos it ive . The OTDR signal suffers the splice loss twice, but

since the scale is typically halved in dBs, the bidirectional

average, equal to (2), is automat ically included.

OTDR Apparent Loss

The OTDR signal level depends on the product of the overall

scattering coefficient and the backscatter capture fraction. For

single-mode fiber, the latter varies as [2] (X/wQ

)2. Hence, if

the mode diameter 2wQ decreases across a splice, compared with its

value for input fiber 1, there will be a corresponding increase in

backscattered signal level in output fiber 2. This means that the

splice loss as seen by the OTDR will appear to reduce by the

backscatter ratio in halved dB's:

B(R) = 5 login( w 2/w l)

2

(3)

- -4.343 R for |R| < 0.35

81
From (2) and (3), the OTDR apparent loss is then

L(R) = dB(R) + B(R)

(4)

= 4.343 (R 2 - R)

Since |r| < 1, the backscatter effect due to mode diameter

mismatch overwhelms the loss effect due to the same cause. Figure

1 schematically shows the result (4) that differs depending upon

the OTDR direction; a positive R will product apparent 'gain'.

Bidirectional averaging naturally produces the best estimate of

real loss:

0.5 [ L( R) + L(-R) ] = dB(R) (5)

Experiment

Gaussian mode diameters were obtained at 1300 nm on both ends

of spooled fibers using the knife-edge technique [3] . Fiber pairs

were purposefully selected to give mode diameter mismatches over a broad' ±16% range, and were spliced via core position monitoring

in an arc fusion machine. The recoated splice was prooftested.

First, OTDR measurements were taken in both directions. The

splice was then cut off the spools, leaving 5 m of fiber on each

side of the splice. The true splice loss was measured by cutback

in each direction. The spooled fibers were respliced and

remeasured several times for the same mode diameter mismatch, and

all readings were recorded in the accumulated data.

Figure 2 shows the correlation between measured bidirectional

averages of cutback loss (which showed little directional

dependence) and OTDR apparent loss (which showed much). Figure 3 displays the directional OTDR apparent splice loss correlated with

mode diameter mismatch; the best fit line to the data is given.

In view of the data scatter common to splice measurements, both

Figures agree reasonably with predicted results. Any small

offsets are likely due to directional imperfections in the splice

procedure, exclusive of mode diameter mismatch.

Conclus ions

The effect of mode diameter mismatch on both splice loss and

OTDR loss has been accounted for. The former can be predicted from

the latter. Eliminating the relative mismatch R between (2) and

(4), or Figures 2 and 3, gives the real loss estimate

dB(R) = (4.343/2) + L - [(4.343/2)2 + 4.343 L ] 1/2 (6)

Care must be taken to include the sign of the OTDR loss L.

82
Finally, a mode diameter tolerance in terms of OTDR

deflection is much tighter than one based on real loss. For

example from (4), an OTDR reading L(R) = 0.3 dB corresponds to a mode diameter mismatch of R = -6.5%. This yields a real power

loss (2) of only dB ( R) = 0.018 dB.

References

[1] For example, W. Anderson and D. Philen, J. Lightwave Tech.

LT-1 20 (March, 1983). An older definition of mode field

radius is used.

[2] A. Hartog and M. Gold, J. Lightwave Tech. LT-2 , 76 (April,

1984). The V-value and older 1 NA 1 terms must be simplified.

[3] W. Otten and F. Kapron, OFC '86 , paper TUL30 (February,

1986); also submitted, with T. Olson, to J. Lightwave Tech.

fiber distance

OTDR

Signal

fiber distance

Figure 1. Schematic Bidirectional OTDR Splice Signals L with

Mode Diameter Mismatch (w^ > w

2

) . Shown are the contributions of true loss dB in (2)

and backscatter ratio B in (3) to apparent loss L

in (4) .

83
Average Cutback Loss (dB)

0 . 3

0 .2

0.1 " x X X

x xxx. XXX X

X X

X XXX

X XXX

x x X X XX XX X XXX

1 X 1 1 0.1 0.2 0.3

x — data

^one-to-one

Average OTDR

Loss (dB)

Figure 2. Correlation of Splice Losses

Measured by Cutback and by OTDR.

(Bidirectional averages are taken in each case; see (5)

)

05Q

O

a —

O CQ H T3

+J —

O

0) w

u w H O

Q ^

X

x* x

X «

1 t 1 1 t ( 1

ry^**^^ x

XX

X

1 t 1 1 1 1 1 1

—data

fit

(4)

Mismatch R (%)

-10 10

Figure 3. Directional OTDR Splice Apparent Loss

vs Mode Diameter Mismatch

(best fit: L = 3.935R2 - 3.607R + 0.030)

84
MULTIMODE FIBER SYSTEMS CHARACTERIZATION

LP. Vayshenker, D.R. Hjelme, and A.R. Mickelson

Department of Electrical and Computer Engineering

University of Colorado

Boulder, Colorado 80309

Introduction

When designing a multimode fiber system (MFS) one should take into consideration the

mode-dependent properties of the components of the system. Lack of a design tool for fiber- optics networks makes the problem more difficult. However, one can describe a MFS using

mode dependent matrices for each component of the system and obtain the system response as a product of these matrices and the input vectors [l]. This method has been used by Evers

[2,3] to characterize couplers and connectors, and by Mickelson et al [4,5] for fusion splices

and microbends. To our knowledge there exist no work on transmission matrices for fibers,

but only mode dependent attenuation measurements [6]. The purpose of this paper is to give an accurate measurement technique to obtain the transmission matrix for arbitrary fiber optic

connectors attached by multimode fibers, and apply these techniques to various

connector/ fiber combinat ions.

Theoretical Relations

The mode transmission matrices can be obtained from power loss and near-field measurements. Using the mode continuum approximation [7], one can relate the near field, I(s), to the

modal power distribution p(R). For an azimuthally symmetric fiber with arbitrary index

profile n

2 = n2 [l — 2Af(s)] this relation is I(s) = J* p(R)RdR where s = r/a is the nor17

f"(.) _ malized radial coordinate, R

2 = 77^~"(1 —

P

2/n 2 k2

) is the mode parameter and V = njka^2A.

It should be mentioned that this approximation is valid provided that the linewidth, 8\, of

the exciting source satisfy 8X/X > ^2A/(ka N/), where N, = n, — X ~- ' s tne 'material

d\

group index. The total power, P,, in the modes with R values between Rj and R1+1 are given

Ri+1 by Pj = J p(R) m(R) dR where m(R) = V

2 R[f_1 (R2 )/a] 2

is the modal density. Dividing RR; space into n intervals, one can define a power vector P = [P l P2

• • Pn ] T (T for transpose).

The input, P('), and output, P' 0

', vectors can then be related by the mode transmission matrix

T [l] as P(°) = T P('). The diagonal elements of T represent the fraction of power incident on the system in the modes of one R-interval which remains in the same R-interval, and the

off-diagonal elements represent the fraction of power coupled between the modes.

For a fiber, T would obviously be length dependent. In a fiber with only absorption and

Rayleigh scattering it would be reasonable to assume T to be a diagonal matrix with ele- ments Tjj = e

7 ' Z

, where 7j corresponds to the differential mode attenuation (DMA) coefficient

•y(R) [6]. But as shown in [6]', other mode dependent effects such as mode coupling and coupling to radiation modes gives "y a length dependence and worse, it is extremely hard to determine from measured data. These problems could be overcome if we define a length

85
independent matrix A as T = e ZA for the fiber. A would correspond to a two-variable DMA

coefficient -y(R,R') as suggested in [6].

Measurement of the Transmission Matrix

The transmission matrix contains in general n

2 independent elements and to obtain

these, one needs n

2 independent measured quantities. This could be obtained using selective

excitation to create n independent input vectors Pjj) and measure P[°). Then T is obtained

from T = [Pj°\ . . . ,Pi°'][P|'\ • • • »Pji The diffraction limit poses a severe restriction on how many independent inputs one can generate. Increasing n would then make T unstable,

and there should be an optimal number of excitations, n. Some authors [2,3] recommend n =

3 and others [l] n = 2. In this paper n = 2 is used. It should be pointed out that the resulting matrix T is very sensitive to the power loss ratios |Pj^|/|P[''|. The larger n and hence the

less independent inputs, the more sensitive T is to this ratio.

The measurement set-up is shown in Fig. 1. Selective excitation is generated by varying

the 3 parameters of the launch system. These parameters can be defined by

u = —, p =

ber>rTI

, t = —

be,Tm where r5 is radial launch position, abeam is the beam radius, a a NAfiber NAbc, am is the effective numerical aperture of the launching system. In the experiment

p = .14, t = .68, u = 0, .6-1 were used. Fig. 2 shows the calculated (using the theory of [6]), p(R) and I(s) for these parameters. To find the input vector P('), the measured near-field

intensity of a short fiber, stripped for cladding modes, was used. Similarly for the connector a

short fiber was used at the output end.

Beside errors in the power measurements and the near-field measurements, other error sources can affect the results. Among these are fiber end quality and alignment problems [8] and leaky modes [9, 10, 11]. According to Petermann [11], the leaky mode correction factor

given in [9.10] is very sensitive to non-circular perturbations of the fiber. The correction

therefore becomes very uncertain and in this work it is not used.

Resu Its The following combinations were under test: Long fiber-connector-long fiber (lcl) and

short fiber-long fiber (si). Fig. 3 illustrates the measured near-fields and derived modal power

distributions for the si combination. To reduce the noise in the data, the measured nearfields were filtered using a fast-fourier transform and a properly chosen cut-off frequency.

Next the filtered near-fields were fitted to a 12th order polynomial (only even powers of s) and

p(R) was calculated and the power vectors formed. The transmission matrix for the fiber is found from the si combination to be

AL [—'.09 "#] <*>

The total transmission matrix, T| c) , for the lcl combination were found to be

rp [21 .051 ^1 "

L.04 .12j

Using this matrix and the matrix for the fiber, the connector transmission matrix can be

86
found using the relation T|c

| = e AL Tc e AL . The resulting connector matrix becomes

T = |\84 .Oil

ic

L.04 .48j

This matrix is in good agreement with the matrices found by Evers [3]. From the fiber

transmission matrix A it is possible to find the steady-state distribution. We can write

p(o) — e -AL p(i) = 2 M-k ' P^1 ' M-k e

k "when jjl k and X k are the eigenvectors and eigenvalues

n=l

of A. After a long propagation distance only the vector corresponding to the smallest eigenvalue will survive, i.e., it is the steady state distribution [12]. The matrix (*) has the following eigenvalues and eigenvectors

h<i =

p.f] ; ^2 -

[_V°o] LX i = -*i> LX 2 = .83.

The eigenvalue \ l corresponds to a steady state loss of 2.63 dB/km. The corresponding power

vector jjLj, is in agreement with the steady state modal distributions found in [4,5].

Acknowledgment

This work is funded by a grant from AMP, Inc., Harrisburg, PA.

References

[I] G.T. Holmes, "Estimation of concatenated system response based on measured transfer

functions for low and high order modes," in Prod. Seventh European Conf.'on Opt.

Commun (Copenhagen, Denmark), Sept. 1981, pp. 3.4-1 - 3.4-4.

[2] G. Evers, "Measurement of mode transition matrices of quasi-step-index optical fiber

components," SPIE vol. 500, Fiber Optics: Short-Haul and Long-Haul measurements and

application II, 1984.

[3] G. Evers, "Mode transition matrices for fiber-optic connectors," El. Lett. 21, 401-2, 1985.

[4] A.R. Mickelson, M. Eriksrud, S. Aamlid, N. Ryen, "Role of the Fusion Splice in the Concatenation Problem; J. Light Technology, Vol. LT-2, No. 2, pp. 126-138, 1984.

[5] A.R. Mickelson, "Backscatter Readout from Serial Microbending Sensors," J. Light

Tech., Vol. LT-2, pp. 700-09, 1984.

[6] A.R. Mickelson and M. Eriksrud, "Mode-dependent attenuation in optical fibers," J. Opt.

Soc. Am., vol. 73, No. 10, pp. 1282-1290, 1983.

[7] A.R. Mickelson and M. Eriksrud, "Mode-continuum approximation in optical fibers,"

Opt. Lett., vol. 7, pp. 572-574, 1982.

[8] D. Rittich, "Practicability of Determining the Modal Power Distribution by measured

Near and Far Fields," J. Light Technology, Vol. LT-3, No. 3, pp. 652-661, 1985.

[9] M.J. Adams, "Correction factors for the determination of optical-fiber refractive-index

profiles by the near-field scanning technique," El. Lett. 12, pp. 281-83, 1976.

[10] M.J. Adams, "Length-dependence effects due to leaky modes on multimode graded-index

optical fibers", Opt. Commun. 17, pp. 204-09, 1976.

[II] K. Petermann, "Uncertainties of leaky mode correction for near-square-low optical

fibers," El. Lett. 18, pp. 513-14, 1977

[12] D. Marcuse, Theory of Dielectric Optical Waveguides, Academic Press, 1974.

87
SPOTSIZE

PINHOLE

HP

216 Fig.l Near-field measurement set-up

Fig. 2 Theoretical I(s)-a and p(R)-b for Fig. 3 Measured I(s)-a and p(R)-b

the excitions used in the eperiment f or ^ ne i on g fiber

88
Single-Ended Bandwidth Measurement for Multimode Optical Fiber Using Far

End Pulse Reflection

Shin-ichi Furukawa, Katsuya Yamashita, Mitsuhiro Tateda

and Yahei Koyamada

NTT Electrical Communications Laboratories

Tokai, Ibaraki, 319-11, Japan

ABSTRACT

Single-ended bandwidth measurement using optical pulses reflected

at the far end of a multimode optical fiber has been realized. This

method makes it possible to measure the bandwidth of the optical fiber

under test with measurement equipment located only at one end of the

test fiber. Consequently, the time and manpower required for bandwidth

measurement of installed fibers are extremely reduced.

1

.

Introduction

Usually, optical fiber bandwidth measurement requires to locate

transmitting and receiving equipment at opposite ends of the fiber under

test^ ^ . Recently, a remote-ended time domain bandwidth measurement

technique has been proposed where both ends of the fiber can be

separated by more than 10 However, the situation where the

transmitting and receiving equipment are located at both ends has not

been changed. In the case of subscriber transmission lines, the

subscriber ends of the fibers are distributed throughout many locations

in the subscriber area, whereas the other ends are centered at the

telephone office. If transmitting and receiving equipment need to be

located at both ends, as usual, a large amount of time and manpower are required for bandwidth measurement. This paper describes a new

bandwidth measurement method in which the transmitter and receiver are located at the same end of the test fiber, thus significantly reducing

time and. manpower.

2. Single-ended bandwidth measurement

The configuration of the single-ended bandwidth measurement is

shown in Fig. 1. The test fiber ends are substantially clean, smooth

and perpendicular to the fiber axis, so that Fresnel reflections at both

ends can be expected. A train of nanosecond pulses is launched into the

fiber under test through an optical directional coupler. The backward

89
pulses reflected from the far end of the test fiber are. detected with an APD receiver through the optical directional coupler (ODC) . Averaging

of the received pulse waveforms is made with a sampling oscilloscope and

transformation from the pulse waveform to baseband frequency response is

accomplished with a mini-computer. The roundtrip return bandwidth BWr

of the test fiber is calculated from the ratio of the baseband frequencyresponse obtained from the far end reflection to the one obtained from

the input end reflection. The transmission bandwidth BWt of the test

fiber is estimated from the roundtrip return bandwidth using the

equation

mt = BWr x 2* , (1)

where V is a correction coefficient for a roundtrip return, pulse in the

test fiber. The parameter If is usually used to describe the bandwidth

characteristics of concatenated fibers^.

A diagram of the optical power level is shown in Fig. 2, where the

averaging of the received pulse waveforms in the. sampling oscilloscope

is not performed. 7 dB optical power loss is allowed for the test

fiber. When 15 dB SNR improvement in optical power level can be

obtained by averaging the pulse waveforms, it is possible to measure the

bandwidth of an optica], fiber having a one-way loss of up to 11 dB.

3. Experiment

An example of the baseband frequency response obtained from a

single-ended bandwidth measurement is shown in Fig. 3. The insets in

Fig. 3 represent pulse waveforms reflected from the input end and the

far end of the test fiber. A test fiber about 12 km long was prepared

by concatenating 4 pieces of fiber with a total loss of 8.5 dB. The

peak optical power of the input pulse in the test, fiber was about 12 dBm

and its full width half maximum (FWHM) pulse width was 1.5 ns. The

wavelength was 1.3 pm. A standard 50/125 urn graded- index multimode 10

meter long optical fiber cord was used to connect the test fiber to the

optical directional coupler. The far end of the test fiber was prepared

using a fiber cutting tool^. The average number of pulse waveforms

in each train was 1024 providing an SNR improvement of about 15 dB in

the optical power level. Deviation from the mean value of the

single-ended bandwidth during 10 measurements under the same condition

was less than 3 %.

4. Discussion

The measured bandwidth changes due to mode conversion when the

cutting angle of the far endface is not perpendicular to the fiber axis.

90
Figure 4 shows the changes in the roundtrip return bandwidth BW

r when the far endfaces of fibers about 2 km long were polished at severalangles. It was confirmed that the change in BWr was less than 7 %

when the angle variation was less than 2°. Usually, the cutting angle

is as small as 0.6°^. Therefore, the influence of the acting angle

variation is negligibly small.

Figure 5 shows a comparison of the transmission bandwidth BWt

estimated from Eq. (1) with the measured transmission bandwidth for

several test fibers having lengths from 1.5 km to 12 km. The

transmission bandwidth measurement was performed by locating the

receiving equipment at the far end of the test fiber without changing

the launching condition. A value of 0.8^ was used for the

correction coefficient X which was determined from several experimental

results. Furthermore, it was experimentally confirmed that If did not

depend on the launching condition. The average difference between the

estimated value and the measured value was as small as 9.2 % and the

maximum was 19.5 %. This accuracy is considered to be sufficient for

the bandwidth evaluation of fiber installed in the field.

5. Conclusion

A single-ended bandwidth measurement is proposed in which a transmitter and a receiver are located at the same end of a test fiber.

With this method it is not necessary for technicians to go to or to

carry measurement equipment to the far end of the test fiber.

Consequently, a significant reduction in time and manpower can be

achieved. By combining this measurement method with the conventional

optical time domain reflectometer , attenuation and bandwidth, which are

the two essential transmission characteristics in multimode optical

fiber, can be measured at one end of the test fiber.

The authors would like to express their grateful appreciation to

Fumihiro Nibei, Naoya Uchida and Nobuya Kojima for their encouragement.

References

(1) CCITT Recommendation ^'Characteristics of 50/125 urn multimode graded

index optical fiber cables", Vol. DL , Fascicle 1.2, Rec. G. 651 (1985)

(2) A. Ahmad et al. ; "Remote-ended time domain bandwidth measurements on installed regenerator sections of multimode optical fiber cable in

the british telecom network", Proceedings of the 32th TvJCS, p. 4 14

(1983)

(3) M. Matsumoto et al. ;"A simple and practical cutting tool for

optical fiber end preparation", Trans. IECE Japan, E66, 11, p. 661

(Nov. 1983)

91
OPTICAL

DIRECTIONAL COUPLER

PULSE GENERATOR

PULSE LASER T

VARIABLE DELAY SAMPLING OSCILLOSCOPE

FIBER UNDER TEST 0

APD

RECEIVER

MINI COMPUTER

Fig. 1 Configuration of single-ended

bandwidth measurement.

Pulse waveform

reflected from

input end of the

fiber Pulse waveform

reflected from

far end of the

fiber

200 400

Frequency, MHz

Fig. 3 An example of the baseband

frequency response obtained

from single-ended measurement

(fiber length = 12 km).

1000

1

> 500 -

A

-o

raE

ID MX

[TRANSMITTER] I QDC I I TEST FIbIr]

0 500 1000

Measured bandwidth, MHz

Fig.

Em

T3

soa

oaO

10

-10

•20

-30

-40

LD output f power

Coupling loss

Fresnel

reflection loss

Pulse peak power down due

to dispersion Y r r 7 r 77 7 7777;

> Fiber loss '

Coupling loss

- Measurement SNR margin

-APD minimum

received power RECEIVER

Fig. 2 Optical power level diagram

for single-ended bandwidth

measurement . Averaging

of the received pulse

waveforms is not performed.

ID output power is 12dBm.

APD mmimum received power

is -38dBm.

(DW

C

1.5 -

o No.1 fiber a No.2

No.3 a No.4

2 4 Angle variation, degree

Fig. A BWr change with angle

variation of the far endface.

5 Comparison of the estimated

transmission bandwidth BW*. with

measured transmission bandwidth

92
EFFECTIVE FIBER BANDWIDTHS

IN LED BASED SYSTEMS

Dieter Schicketanz and William S. Jackman

SIECOR Corporation, Hickory, N.C., USA, 28603

INTRODUCTION

LED's create special problems in optical fiber systems when trying

to predict system performance. With LED's operating close to the zero dispersion wavelength of the fiber system, not only are equations to

predict fiber and system bandwidth complex, but measured bandwidths

using LED's as sources are influenced by the inherent bandwidth

limitations of LED's and are not easily used for system performance

predictions

.

Most predictive equations assume gaussian rolloffs and/or impulse

responses to give baud to bandwidth ratios of 1.2 to 1.6 for specific

power penalties and bit error ratios. \1\ While there is normally a "safety factor" built into bandwidth specifications to account for the

non-gaussianity of modal bandwidth, the extreme non-gaussianity of the

material component of fibers with LED sources operated near the zero dispersion wavelength overwhelm any reasonable "safety factor".

This paper uses the RMS pulse width of the output pulse to

calculate an "effective" bandwidth to satisfactorily predict system

performance with normal approximate gaussian assumptions. Since the

"effective" bandwidth is not directly measurable, system rise times have

been tested and compared to the predicted value indicated by the

"effective" bandwidth. Excellent agreement has been observed.

THEORY

Basic assumptions made in the calculations are: 1) The LED spectrum

can be approximated by a gaussian [S(A.)] with the same "middle"

wavelength [A. ] and same RMS spectral width [a.] as the real spectrum, m A

And 2) The dispersion characteristics [M(A)] of the fiber can be

adequately defined by a Sellmeier equation with a zero dispersion

wavelength [A.^] and a slope at the zero dispersion wavelength [S^] . The impulse response [h(t)] of the fiber to an LED source can be

derived, but is very involved. The RMS pulse broadening [a^] can be

approximated by \2,3,4\:

93
(1) a

t=L[M(A) 2 a

A

2+ (S^V/2] * 5

When A =A. , h(t) can be approximated by:

(2) h(t) = (7ttS

0

La

A

2 )"- 5 exp[-t/(S

0

La

A

2

)]

with a approximately equal to:

(3) Q

t

= LS

Q a

A

2

/ (2)' 5

. When !^- c

~ Aq j >>^ , h(t) can be approximated by a gaussian:

(4) h(t) = exp(-.5{(t-t

0 )/[M(A)La

A

]} 2

) with approximately equal to:

(5) a

t

= LM(A)a

A

. Fourier transforms can be used to predict the frequency response

rolloffs [H(f)] and the 3 dB optical bandwidths [BW]

.

For A = A. the transform of (2) is \5\: m 0 >

N (6) H(f) = (l+j27tfLS

0 a

A

with magnitude:

(7) |H(f)| = (l+(2nfLS

0aA

2

) 2 )"' 25

3 dB optical (jH(f) | = .5)

(8) BW(A

m =A0

) = (15)- 5 /(27iLS

0aA

2

) = .6l6/(LS

0 a

A

2

) = .436/a

t For | A -A. | >>a. the transform of (6) is: m 0 A

(9) H(f) = exp(-.5(f/crjIt)=exp (- . 5 (f 2ttq r)=exp (- . 5 (f 27iLM(A)a, ) ) A

3 dB optical for |A -\n \ » o. m 0 A

(10) BW(|Am

-A.l » a.) =(ln(4))- 5

/(27ta ) =.187/a =. 187/ (LM(A)a. ) m U A t t A

The effective material bandwidth is determined by using the rms pulse spreading found in (1) with the gaussian assumptions which are the

basis of (10). Therefore the effective material bandwidth is:

(11) BW(effective) = . 187/ (L(M(A)

2Q

2

+S„ V 4 /2)

5

) A 0 A

The system bandwidth is a combination of the modal, material and

LED responses. The measured bandwidth [BW(measured) ] using an LED source

can be predicted by multiplying H(f) [modal] by H(f) [material ] and

H(f)[LED] and taking the half power point. H(f) [Measured] is

non-gaussian and depending on the rolloff characteristics of the LED's,

may not be comparable between LED's with the same optical

characteristics . The effective system bandwidth can be calculated using common gaussian assumptions as:

94
2 2 2

(12) l/BW(system effective) =l/BW(modal) +l/BW(ef f ective material)

Limits to the accuracy of the effective system bandwidth occur when

the system is long enough that spectral filtering significantly distorts

the spectral output of the fiber or for extremely high performance or non-gaussian LED's. Spectral filtering could be accommodated if the

spectral attenuation signature of the fibers to be used are sufficiently

characterized and consistent. An "effective" middle wavelength and RMS

pulse spreading could be calculated.

Because the effective system bandwidth is difficult to measure, to

test the theory, rise times of systems were measured and compared to the

predicted rise times using gaussian assumptions using the RMS pulse

width, the same basic assumption behind the effective material bandwidth

calculation.

Rise time calculations from the RMS pulse width start with a normalized gaussian impulse.

(13) h(t)=exp(-.5(t/a

t

) 2

The impulse rise time from h(t) = . 1 to h(t) = .9 is:

(14) r t

. = ((-21n(.l))

,5 -((-21n(.9))' 5 )a

t = 1.6869 a

t

The step response is related to the impulse response:

(19) Step response h(t)= 1/ (27tat

)

5

_ ooJ

t

exp (- (t-t

Q

) 2/ (2cJ

t

2

)dt

The step rise time from h(t) = . 1 to h(t) = .9 is:

(19) r = 2.563o

ts t

Measurements are easier to do for step rise times than impulse rise

times. To measure the rise time of the fiber, it is necessary to

deconvolve the effect of the transmitter and receiver from the actual

measured rise time. Using Gaussian assumptions the system rise time is

related to the component rise times by:

(22) rt

2 = I rt . 2 system components

TEST PROCEDURE AND RESULTS

A Tautron 325 BER test set was used to drive a high performance

LED , k =1290nm and AA=135nm. A Ge avalanche diode (r <200ps) , 5 GHz m ts

amplifier, and a sampling oscilloscope completed the test set up. The

system was first tested without a fiber to measure the test set up rise

time. As shown in figure 1, The non fiber portions have a step rise time

of 2.7 nanoseconds. Figure 2 includes a 2.3 km, 62.5 |Jm core, .28 NA

2

fiber, A

Q

=1345nm, S Q

=.1 ps/nm km, with a modal bandwidth of 247 MHz

95
(800 MHz. km). Expected modal step rise time = 1.4 ns . The predicted

effective material bandwidth is 198 MHz, and the predicted material step

rise time =2.4l3ns. The predicted total system step rise time was 3.9

nanoseconds. Figure 2 shows a measured step rise time of 4 nanoseconds.

Literature

:

\1\ Personic, S. D.; "Reciever Design for Digital Fiber Optic

Communications Systems, I"; The Bell System Technical Journal, Vol.

52, No 6; July-August, 1973; pages 843-875

\2\ Marcuse, Dietrich and Lin Chinlon; "Low Dispersion Single-Mode

Fiber Transmission- The Question of Practical versus Theoretical

Maximum Transmission Bandwidth"; IEEE Journal of Quantum

Electronics, Vol. QE-17, No. 6; June 1981; pages 869-878

\3\ Geckler, S.: "Lichtwellenleiter fur die optische

Nachrichteniibertragen" ; Berlin; Springer 1986, Chapter 3

\4\ Gloge, D. "Effect of Chromatic Dispersion on Pulses of Arbitrary

Coherence", Elec. Letters 21, 1979, pages 686-687

\5\ Cohen, Leonard G. , Mammel, Wanda L. , and Lumish, Stan; "Dispersion

and Bandwidth Spectra in Single-Mode Fibers"; IEEE Journal of

Quantum Electronics, Vol QE 17, No 1, pages 49-53

LED rise time = 2.7 ns LED + Fiber rise time = 4.0 ns Figure 1 Figure 2 THEORETICAL = 3.9 ns

96
MEASURING 825 NM TRANSMISSION ON 1300 NM SINGLE-MODE FIBERS

Catherine M. Ragdale* and Felix P. Kapron

ITT Electro-Optical Products Division

7635 Plantation Road

Roanoke, VA 24019-0065

Abstract

Standard single-mode fibers, designed for longwave

transmission, are operated in the shortwave region well below the

effective cutoff of the second-order mode group. The temporal

and frequency responses are investigated.

Int roduct ion

The advantages of single-mode transmission are realized if

the source wavelength is higher than the effective cutoff

wavelength [1] of the fiber. This will not occur with 1300 nm

transmission over short fiber lengths for which the second-order

mode group is not fully attenuated, and modal noise can result

[2]. However, here we examine operation at even shorter

wavelengths where that mode group becomes less lossy, and study

825 nm fully bimodal transmission over longer lengths of

single-mode fiber designed for 1300 nm operation.

Second-Order Cutoffs

The fibers used were dispers ion-unshifted (EIA fiber type

IVa) , with a nominal step-index matched-clad profile and a small

central dip. The cutoff wavelength of the LP^ second-order mode

group was obtained by the standard power transmission procedure

[3] for a 14-cm radius single turn and a 2-m length. We extended

this procedure to wavelengths short enough to detect cutoff of

the third-order LP mode group. Moreover, the effective cutoff

wavelengths for 1.5 km lengths were obtained [4].

* currently with Plessey Research, Caswell, Northants, U.K.

97
Figure 1 shows results for one particular fiber. For a simple ideal profile, the theoretical cutoff wavelength of the

third-order mode group should be (2.405/3.832) = 62.8% of the

theoretical second-order mode group cutoff. Note that the

measured third-order/second-order effective cutoff ratios are

68.6% and 68.4% for the short and long fibers, respectively, in

reasonable agreement. Moreover, the long-length/short-length

effective cutoff ratios are consistently 94.9% and 94.6% for the

second-order and third-order mode groups, respectively.

The farfield patterns of Figure 2 were made with two light

injection conditions. They show that at 825 nm, both mode groups

are tightly bound over the full fiber length. The third-order

group, which has a long-length effective cutoff of 836 nm, is

comparatively weak.

Bimodal Distortion

A time-domain measurement, done similarly to pulse

broadening with multimode fiber, is shown in Figure 3.

Differential mode delay is responsible for the pulse separation

Ax = 3.65 ns/km (assuming length linearity and negligible mode

coupling). This agrees with the observed frequency periodicity

Af = 1/At = 275 MHz-km in the Fourier-transformed baseband

response A( f )

• cos ( 2tt f LA t ) .

We associate the larger and relatively delayed output pulse

with the second-order and weakly guided third-order mode groups.

This is consistent with the relative mode power degeneracy and

with theoretical step-index group delays [5]. Chromatic

dispersion causes the broadened pulses of Figure 3b compared to

the input width in Figure 3a. It also causes the general

frequency rolloff A( f ) of Figure 3c. The extra broadening of the

larger pulse is due to the weakly guided third-order mode group.

This was confirmed by wrapping the fiber onto a smaller 13-cm

diameter spool; this equalized the widths (but not heights) of

both output pulses.

Conclus ion

Shortwave transmission in a fiber designed to be single-mode

in the longwave region has been studied. Results are consistent

with long-length bimodal operation around third-order mode

cutoff. Such transmission can be useful for some bidirection WDM

high/low bitrate applications.

98
References

[1] H. T. Nijnuis et al, Symp. O.F. Meas., NBS Special Pub. 683 , 11 (Oct 1984)

[2] D. G. Duff et al, paper TUOl , Tech. Digest O.F. Comm. (Feb

1985), and references therein

[3] EIA FOTP-80, CCITT Recomment at ion G.652

[4] EIA FOTP-170

[5] L. G. Cohen et al , B. S.T.J. J39", 1061 (July-Aug 1980)

POWER

RATIO

(dB)

i i r I i i i i l 800 1000 1200 1400 1600

WAVELENGTH (nm) 2 ,s, „o 2

Figure 1. Measurement of short-length and long-length effective

cutoff wavelengths for the second-order and thirdorder mode groups. Power transmitted through the

single-mode fiber is ratioed to power through 2 m

of multimode fiber.

99
Figure 2. Far-field patterns obtained by butting the 1.5 km

test fiber to a similar fiber pigtailed to an 825 nm

laser diode: (a) injection of all modes; (b) transverse

offset injection of mainly the second-order and thirdorder mode groups . The dashed curves were taken with

20 turns around a mandrel of 15-mm diameter.

A. INPUT PULSE

Figure 3. Temporal response of

fully bimodal transmission at

825 nm, along with the Fouriertransformed frequency response.

t(ns)

100
CHARACTERIZATION OF SOURCES AND DETECTORS FOR OPTICAL

FIBER SYSTEMS

T.ITO* and K . KURUMADA* *

NTT ELECTRICAL COMMUNICATION LABORATORIES

*l-2356 Take, Yokosuka-shi , Japan

** 3-1 Wakamiya, Morino-sato, Atsugi-shi, Japan

1. INTRODUCTION

Fiber optics are now playing an important role in

telecommunication networks that include inter-city trunks,

undersea transmission lines and subscriber loops. Present

fiber optic communication systems utilize an

intensity-modulation direct-detection scheme. However,

advanced lightwave technology using coherent light is

expected to go beyond the limit of present fiber optic

communication systems. Optical device characterization is

key technology in the development of new fiber optic

communication systems.

This papers describes the present state of semiconductor

light source and photodetector characterization

technologies, while reviewing the progress made in the

development of fiber optic communication systems at NTT.

2. TRUNK TRANSMISSION

400 Mbit/s single-mode fiber trunk lines are now

connecting major cities in Japan. The following items are

important in designing high-speed long haul transmission

systems

2.1 LP transversal mode

Stable single transversal mode oscillation is

indispensable to obtain large a signal-to noise ratio. The

LD transversal mode stability is estimated from a kink of LD

I-L characteristics, i.e. a multi-mode LD has a kink in the

I-L characteristic curve.

The mode stability and single mode oscillation

characteristics are closely related to LD structures.

Among the LD structures proposed thus far, especially the

index-guide LDs including a buried hetrostructure (BH) LD

have shown excellent mode stability.

2.2 LD spectrum

A multi-longitudinal mode Fabry-Perot LD restricts

transmission distance by signal waveform degradation due to

101
the mode partition noise. The degradation is evaluated from

an LD spectrum shape. The LD spectrum, changing with a LD

structure and an operating condition, is not unique. One is

a single-peak shape ; another a double-peak shape. In

general it can be said that, for Gaussian shape spectrum,

the transmission distance limited by this degradation is

given as :

Lm = 350 / w m f/j (km) , where w is an LD average spectrum width (nm) , fg the

transmission bit rate (Gbit/s), and m the single-mode fiber

dispersion ( psec/km nm) (1)

.

2.3 Longitudinal submode suppression

A single transversal mode DFB LD solves the mode

partition noise degradation. In this case, wavelength

(optical frequency) chirping inherent in a DFB LD modulation

restricts the transmission distance. A streak camera, then,

is a powerful tool to examine the dynamic chirping

behaviour. The transmission length restricted by the

frequency chirping is given as :

Lm = 120 / AA m f

Q (km) , where AA is an average spectrum width of frequency

chirping (nm) , fg the transmission bit rate (Gbit/s), and m

the single-mode fiber dispersion ( psec/km nm) (2).

The modulated LD characteristics are evaluated from the

submode suppression ratio. A suppression ratio of more

than 30 dB is required to eliminate the submode effect on

transmission characteristics.

Some sophisticated DFB LDs proposed thus far have

already met this requirement. An asymmetric laser cavity

structure is a promising way to suppress the submode. This

structure is realized by forming an anti-reflection (AR)

coating on one face of a DFB LD chip. Since a low

reflectivity face provides more output power than the other

face, light from the coated face is usually coupled into the

transmission fiber. However, the AR coating significantly

increases the backscattering effect on an LD '

s

characteristics. An optical isolator having more than a 25

dB isolation ratio is able to eliminate the reflected wave

noise

.

2.4 Modulation waveform

The LD modulation waveform, for high speed modulation in

particular, depends on bias and modulation conditions.

102
In a practical transmission system design, the modulated

waveform is specified as masked waveform pattern.

2.5 Reliability

Undersea transmission requires extremely high device

reliability. For an NTT 400 Mbit/s undersea transmission

system (FS-400M system) , Fabry-Perot Buried Heterostructure

(BH) LDs are selected by a threshhold current screening test

and then undergo more than 1500 hour in an aging test to

determine that the LDs have more than a 25 year average

lifetime and a failure rate of less than 50 fit (3)

2.6 APD characteristics

Quantum efficiency, the dark current excess noise

factor, maximum multiplication coefficient, and the

bandwidth are major parameters in system design. An InGaAs

APD, when compared with a Ge APD, displays a low dark

current level, small excess noise factor and longer

wavelength responsibility. Therefore, higher speed

transmission systems such as a 1.6 Gbit/sec system (F-1.6G

system) and 1.5 urn systems use InGaAs APDs.

3. COHERENT TRANSMISSION TECHNOLOGY

Coherent transmission, accompanied with control of wave's

fundamental properties such as frequency, phase and

amplitude, aims at full utilization of broad lightwave

frequency band up to several hundred THz . While several

barriers, listed below, stand in the way of this technology,

there have been enthusiastic research into these area for

the past several years.

3.1 Lightsource phase noise

Microwave or millimeter wave single tone oscillator

spectrum width (phase noise) is less than 1 Hz and , in

practice, the oscillator phase noise does not affect

transmission characteristics. However, in a lightwave

region, it is difficult to obtain the same absolute spectrum

width as in a radio wave region, since the lightwave

frequency is so high that oscillator requirement becomes

relatively restrictive.

The required spectrum width, which depends on a modulation and demodulation scheme, is between 10 - 10

times transmission speed for heterodyne /homedyne detection.

A wider spectrum width results in an error rate floor

phenomena. The most accurate one among LD spectrum width

measurement tools is a delayed self heterodyne/homodyne

technique with an accuracy of about 1 kHz (4).

103
3.2 Lightsource modulation bandwidth

The lightsource spectrum width can be reduced by

inserting an external cavity. However, the external cavity

reduces an LD 1 s modulation performance, i.e. its FM

characteristics. Heavily doped LDs and multi-quantum well

(MQW) LDs are promising lightsources for coherent high speed

systems

3.3 Lightsource frequency stability

Although present intensity modulation fiber optic

communication has, compared with radio systems, larger

transmission capacity, the present fiber optic system

frequency utilization efficiency is in the order of 10 - 10.

The improvement of LD frequency stabilization for optical

frequency division multiplexing (optical FDM) will permit

the full use of a broad lightwave frequency band up to

several hundred THz . At present, LD relative frequency

stabilization is less than 1 MHz. However, highly accurate

optical frequency measurements and semiconductor crystal

growth processes for highly frequency controlled LDs are

required to control absolute optical frequency so that more

than one hundred channel optical FDM transmission can be

realized.

3 .

Photodiode bandwidth

Higher speed transmission requires a broader bandwidth

photodiode. For example, a heterodyne /homodyne detection

demands photodetectors 1.5 - 2 times as broad as its

transmission speed. A wideband photodiode frequency

response over 10GHz is evaluated by the beat of two laser

diodes having slightly different frequencies (5)

4. SUBSCRIBER LOOP APPLICATIONS

For subscriber loops, there are multiple choices among

the combination of lightsources, photodetectors and fibers.

Here as an example, multimode fiber analog modulation

systems will be discussed, because either multimode fiber

analog baseband or FM modulation has been a key economic

factor in fiber optic filed trials. In multimode fiber

systems, S/N degradation due to modal noise is a significant

problem. Basically there are two ways to solve this problem

(1) The employment of modulation schemes, such as FM,

squared wave FM, PFM, which can resit the modal noise.

(2) The reduction of lightsource coherency to suppress modal

noise

104
In terms of economy, an LED is one of the promising

solutions in constructing simple and stable systems. With

regard to the LED systems, keen attention has been paid to

the following items.

4.1 LED wavelength-division-multiplexing

At various NTT subscriber system trials, wavelength

division multiplexing (WDM) , .which makes subscriber system

more flexible and more economical, has used 0.78 - 1.55 ura wavelength. The shortest wavelength for practical silica

fiber subscriber loops is 0.78 um.

Since an LED has a broad spectrum width of 30-100 nm,

it is necessary to reduce interchannel optical interference

by optical filters in optical multi/demultiplexers

4.2 LED nonlinear distortion

Analog modulation, in comparison with digital

modulation, demands LED modulation high linearity. LED

modulation characteristics are usually compensated by an LED

driving circuit. The nonlinear distortion caused in a PIN-photodiode or in an APD is much smaller than that in an

LED. The nonlinear distortion affects video signals

remarkably. Video transmission specifications are usually

described in terms of differential gain (DG) and

differential phase (DP) . The requirement for NTSC video

transmission is a DG of less than 10% and a DP of less than

4° (6).

Acknowledgement

The authors would like to express their sincere thanks to

Dr. K. Nosu, Dr. M. Saruwatari and Mr. E. Yoneda for their

fruitful discussion. They also would like to express their

thanks to Dr. H. Kimura , Dr. S. Shimada, Dr. M. Fujimoto and

Dr. T. Ikegami for their encouragement.

References

(1) Y. Kobayashi et al .

, "Non-repeatered submarine optical

transmission terminal equipment," Rev. of ECL, 31, 6,

pp. 844-850 (Nov. 1983)

(2) K. Iwashita et al., "Error rate performance of high-speed

optical fiber transmission systems using

single-longitudinal mode laser diodes,"

Trans. IECE Japan, J67-B, 12, pp. 1415-1422 (Dec. 1984)

(3) Y.Nakano et al., "Screening method for laser diodes with

high reliability," Electron. Lett., 20, 10, pp. 397-398

105
(May 1984)

( 4 ) S . Kawanishi et al ., "Wideband frequency response

measurement of photodetectors using optical heterodyne

detection technique," Electron. Lett., 22, 6, pp. 337-338

(Mar. 1986)

(5) T.0koshi et al., "Novel method for high resolution

measurement of laser output spectrum," Electron. Lett.,

16, 16, pp. 630-631 (July 1980)

( 6 ) J . Yamagaka et al., "Development of fiber optic broadband

interactive distribution network", to be presented at

ISSLS 86, Sep. 29-Oct.3,

1986, Tokyo.

106
CHARACTERISATION OF SPECIALITY FIBRES AND COMPONENTS

D.N.Payne , S.B.Poole*, M.P.Vamham+

/ R.D. Birch*

* Optical Fibre Group, + British Aerospace pic

Southampton University, Six Hills Way,

Southampton, Stevenage,

England. England.

Optical fibres are making ever increasing inroads into areas traditionally

satisfied by older, more established technologies. One of the prime vehicles

for this trend is the proliferation of specially-tailored fibres designed for

specific applications . Since these applications are diverse (especially in

the sensor field ) and the required fibre characteristics often subtle,

measurement of the fibre response frequently presents a challenge. On

occasions it is even necessary to resort to indirect measurements by

observing the performance of the fibre in the intended application, from

which its characteristics can be implied. A particular example here is the

fibre gyroscope, where signal processing techniques have progressed to such

an extent that the gyro itself provides the ideal environment for

high-accuracy measurements of fibre polarisation performance.

It is the intention of this paper to review some of the many speciality

fibres currently available or being developed, the selection being made to

emphasise topical measurement issues. A summary of different fibre types,

their applications and associated measurement problems is given in Table 1.

Dispersion-Tailored Fibres were among the first speciality fibres to be

investigated and a number of techniques have been developed to measure the

chromatic dispersion. Despite this, questions remain concerning which

curve-fitting routines should be used to reduce measurement noise. This is

particularly true for dispersion flattened fibres, since no a priori

knowledge of the shape of the dispersion curve can be assumed.

2 3

Polarisation-Mairitaining Fibres: ' Controversy exists as to whether

4

these should be characterised by the polarisation holding h-parameter (which

is the polarisation cross-talk per unit length), or by the intrinsic

modal-birefringence B. The h-parameter measures the actual cross-talk

present when the fibre is subject to uncontrolled random external

perturbations, for example by winding on a drum. It therefore reveals more

about the quality of the winding or fibre coating than it does about the

107
fibre. The modal birefringence, on the other hand, describes the resistance

of the fibre polarisation state to cross coupling (externally induced in all

but the worst fibres) and is therefore an intrinsic fibre property. The

problem is complicated by the fact that the modal fields are not plane

5 polarised , but are actually curved, with the orthogonally-polarised

minor-field components present at levels ranging from -32 to -45 dB. Thus

-5 with state-of-the-art fibres providing h-parameters ranging from 10 to

10 ^m ^ (cross-talk of -20 to -30dB after 1 km), it is necessary to use long

lengths of fibre in order to obtain a reliable measurement of h. In this

case, the measurement is usually dominated by external perturbations.

Since the modal birefringence is an intrinsic fibre property, it would appear

to be a better measure of fibre characteristics than the h-parameter^.

Perhaps the best solution, however, would be a measurement of h-parameter

with the fibre subject to controlled perturbations, such as mandrel bends and

twists. In this way the fibre resistance to cross-talk could be directly

specified, since the measurement would automatically include the resistance

of the fibre and buffer coating to cross-talk caused by microbends and

pressure.

7 8 Single Polarisation Fibres ' rely on the differential-mode attenuation

which occurs in depressed-cladding, stress-birefringent fibres when operated

at low V-values ( V~l . 8 ) . Figure 1 shows the spectral differential

attenuation for the two orthogonally-polarised modes in a conventional

Bow-Tie fibre and a polarising Bow-Tie fibre having the same beat length. It

is clear that the polarising fibre has sharper microbending edges and a wider

polarising bandwidth. To obtain reliable measurements, care must be taken to

reduce the effects of stray light and polarisation dependence within the

monochromator

.

The performance of a polariser can only be fully specified by including

cross-talk terms in the form of an intensity transfer matrix, I, between the

g input and output polarised modes . This matrix has been measured for a coil

polariser with broadband light from a laser diode, and the results are as follows:

108
where T^, T are the transmission ratios of the x and y polarised modes and

T T are the cross-coupled power between the modes. First T = -2 dB is xy yx XX

measured using a cut-back technique. X-polarised light is then launched and

the output extinction ratio = T^/T^ = -42 dB measured with a prism

polariser. Similarly, the output extinction ratio FL, = T^/T^, transmission

ratio Rg = (T^ + T VfT^ + T ) = -42 dB is measured using the prism

polariser at the input.

The measurement of T in the above may have been limited by the minor-field

5 yx J 1

components . This is not a limitation for the other terms because the source has the wrong spatial symmetry to launch a mode via its minor-field

components. A similar principle can be used to advantage in mode

extinction-ratio measurements, since a conventional monomode fibre will

spatially filter (to 50 - 60 dB) the minor-field components which pass

through a prism analyser. Following the analyser with a short length of

monomode fibre reduces the minor field components to -80 to -lOOdB. Thus the

measurement is now limited by the prism polariser (typically 50-60dB), not

the minor field components.

Circularly Birefringent Fibres"^ (Fig. 2) are fabricated by rapidly

spinning a preform containing an offset core. A nondestructive method to

measure the polarisation beat length of such a helical-core fibre is to

launch linearly-polarised light and to cross an analyser with respect to the

emerging linear polarisation. The birefringence of the fibre is then

perturbed by moving a pressure-point along its length and counting the nulls

in the cross-coupled power. These occur when the polarisation vector is

parallel or perpendicular to the direction of the force. Thus if N nulls are counted in length L, then L =2L/N. A similar method has been used to measure

P 11

Lp in a linearly-birefringent fibre . A novel feature of the helical-core fibre, is that the fibre is effectively

monomode at high V-values. This is because the higher-order modes cannot

survive the fibre curvature. A simple demonstration of the effect can be

seen in a fibre where the spin pitch has been gradually reduced along its

length. When light is launched into the fibre, bright bands appear along its

length showing where successively lower-order modes are reaching their

effective cut-offs.

Metal/Glass Composite Structures are formed by introducing metal into

109
longitudinal holes close to the core of a silica optical fibre (Table 1).

The fibres make ideal low-loss polarisers utilising the absorption of one of

12 the polarisation states in the metal layer . The measurement requirements

are similar to those for coiled fibre polarisers, but with reduced errors from minor-field components. However, owing to their very high extinction

ratio (>52dB), the optical quality of the crystal analyser and dynamic range

of the spectral loss measurement system become the limiting factors. Further

improvements in measurement dynamic range may be achieved using a gyroscope

configuration^

.

By incorporating two metal electrodes on either side of the fibre core a

14 fibre modulator can be constructed based on the Kerr effect . Here care must be taken in separating the phase modulation produced by the Kerr effect

from that caused by electrostriction.

Rare-Earth Doped-Fibres containing small (<0.25%) quantities of rare-earth

3+ 15 impurity ions (e.g. Nd ) in the fibre core are of great interest as fibre

16 lasers and amplifiers . To this end, a knowledge of their absorption and

fluorescence characteristics is required. Measurement of the absorption

spectra of these fibres poses a number of problems owing to the difficulty of

measuring losses which vary by several orders of magnitude across the

spectrum (Fig.3). However, a multiple cut-back technique in combination with

attenuation equipment optimised for dynamic range (>48dB in single-mode

fibres with 2nm spectral resolution c.f. 35dB in most systems) may be used to overcome these difficulties. Various fibre lengths must be measured,

depending upon the regions of the spectrum under test and the results may

then be re-scaled and combined to give absorption spectra with extremely high

dynamic ranges. Care must also be taken to avoid fluorescence effects.

The fluorescence spectra of the fibres may itself be measured using similar

equipment, but with the fluorescence excited by pumping the fibre with a

laser matched to one of the impurity absorption bands. Examples of

3+ 3+

fluorescence spectra for a fibre containing 450ppm Tb , 30ppm Nd and lOppm

3+

Er ions are shown in Fig. 4.

Fibre lasers present a number of measurement problems in their own right.

Apart from threshold, gain and output power, the spectral output

characteristics are of particular interest for potential applications such as

a broad-band source for the fibre gyro, or as a narrowband optical

communications source. Measurements are complicated by the fact that the

110
axial mode spacing may be only a few MHz, owing to the length of the laser

cavity (up to 300m). Results on Nd3+

(0.9 and 1.08 yum), Pr

3+

(1.07 yum) and

Er

3+

(1.55 yum) will be presented.

Acknowledgements : The authors would like to thank past and present members

of the Optical Fibre Group in the University of Southampton who have made and

are continuing to make significant contributions in this field. Particular

thanks are due to L.S.Li. A Readership (DNP) and a Fellowship (SBP) were

provided by Pirelli General pic. Some of this work was supported under the

DTI/J0ERS programme.

References:

1. D.N.Payne, Proc. OFS, pp. 353-360, Stuttgart, 1984.

2. R.H. Stolen et al., Appl. Phys. Lett., 33, pp. 699-701, 1978.

3. R. D. Birch et al., Electron. Lett., 18, pp. 1036-1038, 1982.

4. S.C.Rashleigh, J. Lightwave Technol., LT-1 , pp. 312-331, 1983.

5. M.P.Vamham et al., Electron. Lett., 20, pp. 55-56, 1984.

6. F.Payne et al., Proc. EOOC, Barcelona 1986.

7. J. R. Simpson et al., Proc. OFC, p. 32, New Orleans, 1983.

8. M.P.Vamham et al., Electron. Lett., 19, pp. 246-247, 1983.

9. M.P.Vamham et al., Opt. Lett., 9, pp. 306-308, 1984.

10. M.P.Vamham et al., Proc. EOOC, pp. 135-138, Venice, 1985.

11. N.Chinone and R.Ulrich, Opt. Lett., CITS, pp. 16-18, 1981

12. L.Li et al., Proc. EOOC, Barcelona, 1986.

13. R.A.Bergh et al., Opt. Lett., 5, pp.479-481, 1980.

14. L.Li et al., Electron. Lett, to be published

15. S.B. Poole et al., J. Light. Tech., LT-4, July 1986.

16. L.Reekie et al., ibid. July 1986.

Optimised singlepolarisation fibre

Non-optimised

high — bi fibre

500 600 700 800 900

Wavelength (nm)

1000

Fig.l. Spectral Attenuation

differential loss between Xwider polarising window for

of Bow-Tie fibres showing

and Y-polarised modes. Note

fibre with optimised design.

Ill
Fig-. 2 . Transverse view of helical-core circularly-birefr ingent

f ibre

.

10,000

1000

Loss

(dB/km)

100

500 700 900 1100 1300

Wavelength (nm)

1500

Fig. 3. Spectral attenuation of a single-mode fibre doped with

lOppm Er . Note very large measurement dynamic range required

Fluorescence

(arbitrary

units)

1 Tb

3> Nd

Er

500 700 900 1100 1300 1500

Wavelength (nm)

Fig. 4. Fluorescence spectra of fibres containing Tb^ +

, Nd^ + and Er^ + ions.

112
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113

Measurement Requirements for

Waveguide Electro-Optic Devices

R. C. Alferness

AT&T Bell Laboratories

Holmdel, New Jersey 07733

Abstract

Electro-optic waveguide devices, including modulators, switches

and tunable filters provide important control functions for

future lightwave networks. The technology as well as the

application of these devices to high performance research systems

has made rapid progress in the last several years. Indeed,

waveguide modulators in limited quantities are now available

commercially. In this talk we will discuss the optical and

electro-optic measurements required to support this rapidly

growing and potentially broadly applicable technology.

Introduction

Waveguide electo-optic devices, including switches, modulators

and electrically tunable wavelength filters have been the subject

of research for more than fifteen years. Significant advances

toward making these devices practical components to provide

essential contol functions for lightwave systems and signal

processing have been achieved in the last several years. This

progress, which includes the demonstration of low i n s e r t i o n- 1 o s s , fiber pigtailed high-speed modulators and complex switch arrays,

has been stimulated by the rapid growth in single-mode lightwave

115
systems. The current interest in extending single-mode fiber

into the loop is particularly important. Such applications

require low cost components which is potentially achievable

with mass produced electro-optic waveguide devices using

photolithographic techniques.

Several technologies for electro-optic waveguide devices are

currently being researched. They are the epitaxially grown

waveguides based upon the semiconductor materials GaAs and InP in

which complete monolithic opto-electronic integration is possible

and the more developed titanium diffused lithium niobate

(TirLiNbOg) waveguide technology. Indeed, commercial TirLiNbOmodulators for research applications have been announced by

3 several companies. Measurements -both optical and electrooptical - are required to evaluate and support the manufacture of

this new technology. In this talk we will review the important

device parameters and measurement requirements. For clarity we

consider the titanium diffused lithium niobate technology, but

the measurement concepts apply to the semiconductor based

technologies as well.

Fabrication

The fabrication steps to fabricate TiiLiNbO^ waveguides are

shown in Fig. 1. Briefly, the pho t o 1 i t ho g r a ph i ca 1 1 y defined

titanium device pattern of thickness x and width W is deposited,

either by electron beam, RF sputtering or a resistively heated

evaporator, on a clean, polished crystal. The crystal is then

placed in a diffusion furnace and heated at a temperature, T, for

t hours. After diffusion, an intermediate insulating layer,

116
typically SiC^ , is deposited over the crystal. Control

electrodes - typically aluminum or, for high-speed devices,

gold-are then deposited over the waveguides. Finally, fiber

pigtails are aligned to the waveguides and attached using UV

curable epoxy. Typical parameter values are: W = 6 - lOum, x =

500 - 1000A, T = 1000°C - 1050°C and t = 4 - 10 hours. The metal

diffuses in as an oxide with an error function profilewhere the

characteristic depth depends upon the diffusion temperature and

time. The waveguide substrate index difference, An, is

proportional to the titanium concentration with a peak value that

depends upon the titanium thickness, and the diffusion time and

temperature. The titanium density is also important and may

depend upon the deposition technique.

The important waveguide parameters are the optical mode shape and

size, which is important for fiber-waveguide coupling and switch

voltage, and the propagation loss. For some devices the

effective index may also be of interest.

Electro-Optic Devices

A large family of electro-optically controlled waveguide devices

has been demonstrated. Examples include the switch/modulators,

frequency shifter, tunable filter, polarization filter or controller. Obviously special measurements are required

depending upon device function. However, the most important

devices currently are the Y - branch interf erometric modulator

and the directional coupler switch/modulator, both shown

schematically in Fig. 2. The transmitted or cross-over

efficiency in the two cases, respectively, varies in responsed to

117
applied voltage through the electro-optic effect as shown also in

Fig. 2. The required switch voltage, an important device

parameter, depends upon the material electro-optic coefficient,

wavelength, electrode length, electrode gap and the overlap, T , between the applied electric field and optical mode. All these

parameters are well known or • photolithographically defined

except, T , which depends upon the diffusion process and

electrode alignment. However, for good electrode alignment, V , can be inferred from the optical mode profile. A typical voltage

x length product figure of merit for TirLiNbO^ switch/modulator

is 6 volt • cm for X= 1.3 um.

These devices can be operated at very high speeds by using a traveling-wave electrode. Therefore, the electrical frequency

response must also be measured. For Ti:LiNbO^ traveling-wave

switch/modulator the f igure-of-merit for the 3dB bandwidth Af is

a bandwidth x length product of ^10 GHz • cm. Broad bandwidth

devices are particularly attractive for high bandwidth lightwave

systems. A schematic diagram of the measurement system for

high-speed measurements in both the time and frequency domain is

shown in Fig. 3.

Measurements

A somewhat simplified measurement pyramid that summarizes the

important parameters to be measured from substrate material, to

fabrication, through the chip stage to the final packaged device.

A discussion of current research measurement techniques as well

as the challenges for future manufacture-compatible methods will

be discussed in the talk.^ - ''

118
References

1. For an overview of guided-wave technology, see, for example,

R. C. Alferness, "Guided-Wave Devices for Optical

Communications", IEEE J. of Quantum Electron., QE-1

7

, p. 946

1981; or R. C. Alferness, "Optical Guided-Wave Devices",

Science, to be published.

For applications in lightwave systems see, for example,

S. K. Korotky, et al., "4 Gb/s Transmission Experiment over

117 Km of Optical Fiber Using a TirLiNbOo External

Modulator", IEEE J. Lightwave Tech., LT-3 , pp. 1027-1031,

1985; or L. McCaughan and G. Bogert, "4x4 Ti:LiNb03

Integrated-Optical Crossbar Switch Array, Appl. Phys.

Letts., 47, pp. 348-350, 1985.

3. See, for example, "Lithium Niobate Devices Finally Make It

to Market", W. R. Iversen, Electonics, J. p. 20, January 13,

1986.

Substrate Index Measurement: G. E. Peterson, S. R. Lunt,

R. J. Holmes, Y. S. Kim, "Refractive Index Measurements of

Lithium Niobate Integrated Optical Substrates", Proceedings

of SPIE, vol. 578 , p. 31, 1985.

Optical Mode Profile: C. Y. Chen and S. Wang, "Near-field

and Beam-Waist Position of Semiconductor Laser", Applied

Physics Letters, 37_, p. 257, 1980.

Fiber-Coupled Insertion Loss: R. C. Alferness,

V. Ramaswamy, S. K. Korotky, M. D. Divino and L. L. Buhl,

"Efficient Single-Mode Fiber to Titanium Diffused Lithium

Niobate Waveguide Coupling for X= 1.32um", IEEE J. Quant.

Electron., QE-18 , p. 1807, 1982.

7. High-Frequency Response: S. K. Korotky, G. Eisenstein,

R. C. Alferness, J. J. Veselka, L. L. Buhl, G. T. Harvey,

and P. H. Read, "Fully Connector ized High-Speed Ti:LiNb03

Switch/Modulator", IEEE J. of Lightwave Tech., LT-3 , p. 1,

1985.

119
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120
Practical Upper Limits to Cutoff Wavelength

for Different Single Mode Fiber Designs

L. Wei*, C. Saravanos and R.S. Lowe

Northern Telecom Canada Limited

Optical Systems Division

8 Colonnade Road

Nepean, Ontario, CANADA K2E 7M6

ri-31 Recently, a number of studies 1 have been conducted in an

effort to establish upper limits for the LPj^ cutoff wavelength

of single -mode fibers designed for 1300nm operation. A value of

1350nm has been quoted for a depressed cladding fiber, design

based on a modal noise analysis . To date, little work has

been done to determine how this upper cutoff value may be affected

by fiber design. In this paper, we utilize the measured length

dependence of cutoff wavelength together with applicable modal

noise constraints to determine the upper limits of cutoff

wavelength for both matched and depressed cladding fibers. For

these calculations, we have not used the wavelength dependent

[3] loss of the LR-j^j mode due to the sensitivity of this

measurement to fiber layout condition. However, by relating the

LP 11 moc^e l° ss to the length dependence of cutoff wavelength,

this experimental problem has been avoided.

The attenuation of the LP,, mode can be approximated by the

* ii • 4-- [2] [3]

11

following equation

X - X^JL)

a uL = 19.4 X 10 ~ (1)

where (L) is the cutoff wavelength measured on a length of

fiber L and X is a constant related to the fiber- index profile.

Comparing Eq. (1) with the widely accepted empirical formula for

[4] the length dependence of cutoff wavelength viz:

L2 X

c

(L 2

) = \ c (Li) + C log j- (2)

where C is a constant we can relate C and X if the cutoff

condition of 19.4 dB is assumed:

(3)

C = -10/log X

* On leave from Research Inst, of Telecomm. Transmission, PTT,

Beijing, P.R. China.

121
Therefore, by evaluating C the wavelength dependent attenuation

of the mode can be determined.

Measurements of cutoff wavelength length dependence for different

fiber designs have been carried out employing the transmitted

power method. For matched and depressed cladding fibers,

(referred to as fiber types I and II in Table 1) the parameters

C were -22 and -55nm/decade with corresponding X values of 2.82

and 1.51, respectively. The latter compare with a value of X =

r 4

1

1.43 as calculated from Anderson's C value of -65nm/decade

for a fiber with a more heavily depressed cladding (fiber type III

in Table 1) . A direct measurement of the wavelength dependent

attenuation of the LP, , mode was performed by a technique

described elsewhere ^ and, as shown in Figure 1, good

agreement with the calculated LP^l mode loss is obtained.

The upper limit of cutoff wavelength is conventionally fixed at

the minimum operating system laser wavelength in order to

eliminate propagation of the mode in the fiber. In fact,

from the system point of view, there is no rationale for this

restriction because, providing the modal noise penalty is kept

small enough, operating in the two mode region does not impose a practical limit to system performance. If Eqs . (1) and (3)

are combined, with equal to the standard measurement length

for cutoff wavelength and I^ = L, the following relation is

obtained

:

A^) = XC (L) + C log

a£* 1 (4)

r 2

1

If the modal noise penalty is chosen assuming reasonable

worst case system parameters for a particular bit rate, the loss

parameter can be determined for the shortest unspliced fiber

length L (e.g. 10 meters). Using this value of a n , the cutoff

wavelength (L,) at the factory test length can be calculated

from Eq. 4 if A c (L) is set equal to the minimum operating laser

wavelength, A g

. A C (L

1

) becomes the maximum allowable factory

cutoff wavelength. The variation of maximum cutoff wavelength

with modal noise penalty is shown in Fig. 2. For A? =0.5 dB,

122
X c (2m) for fiber types I, II and III are 1305nm, 1335nm and

1345nm, respectively. Typical results at different mode

partition noise values (K) , corresponding to different bit rates,

are given in Table 1. Experimental confirmations of these results

utilizing a 2 Gbits/s system will be presented at the conference.

In conclusion, it has been found that by using the measured

length dependence of cutoff wavelength together with appropriate

modal, noise considerations, upper limits to the cutoff wavelength

of single -mode fibers as measured on a 2 meter sample are

established. It has been shown that this limit depends strongly

on the particular fiber design. For matched cladding fibers, the

maximum cutoff wavelength is shown to be 30 to 40nm lower than

for depressed cladding fiber in order to meet the same modal

noise penalty.

References

[1] N.K. Cheung and P. Kaiser, ECOC '84, P. 242, 1984.

[2] D.G. Duff, et al., OFC '85, TU01, P. 52, 1985.

[3] F.M. Sears, et al., OFC '86, TUL19, P. 66, 1986.

[4] W.T. Anderson and T.A. Lenahan, IEEE, Journal of Lightwave

Technology, Vol. LT-2, No. 3, P. 238, 1984.

[5] K.A.H. Van Leeuwen and H.T. Nijnuis, Optics Letts., Vol. 9,

No. 6, 1? 84.

Table 1: Maximum cutoff wavelengths for different mode

partition noise values (K)

" K

Fiber Type "—

0.4 0.6 0.8 1.0

I A + = 0. 25%

A " = 0 % 1312 1307 1305 1302

II A +

= 0 . 2 %

A ~ = 0 , 09%

1355 1342 1335 1332

III A +

= 0.2 %

A

~ = 0 . 15%

1365 1350 1345 1340

123
Fig. 1: Comparison of measured and calculated

LP^ mode attenuation. Calculated values

are derived from the length dependence of

cutoff wavelength.

Fig. 2: Variation of upper cutoff wavelength

limit with modal noise penalty for 3 fiber designs.

System parameters used are K = 0.8, N = 1 dB , L = 10m,

and A = 1285nm.

s

124
New Measurement Technique for Measurement

of Microbend Losses in Single Mode Fibers

A. 0. Garg and C. K. Eoll

Siecor Corporation

Introduction

Microbending is very important to the performance of various fiber

optic devices. Until now, however, the authors have not found a repeatable method for measuring microbend losses in single mode

optical fibers. Various techniques have been suggested and are in use

for multimode fibers. Though these techniques give usable results for

multimode fibers, they are not particularly reliable in the single

mode case. One technique is the sand paper test. This method has the

disadvantage that measurements with two different sandpapers of the same nominal grit size can give different results. Similarly the grit

is not necessarily uniform over the whole sandpaper, and this

introduces further randomness in the measurement. We have found a new measurement technique which avoids much of this uncertainty, and which

gives results which are more repeatable. One can use theory to

predict results of this new measurement technique. We use ASTM

standard sieves of various sizes to give different types and amounts

of microbending. This report describes the methodology and also

presents some of the results obtained using the new technique.

Methodology & Theory

One form of microbending in single mode fibers is caused by

deformations of a fiber in the range from 0.1 to 0.3 mm in length.

Microbend losses in single mode fibers due to this kind of

perturbation show up as an increase in attenuation across a wide band

of wavelengths; the extra attenuation at 1300 nm, for example, can be

comparable to that at 1550 nm. Increasing the length of the

125
perturbations beyond 1mm results in the familiar microbending edge

similar to a bending edge, where losses are very wavelength dependent

i.e. where losses increase dramatically as the wavelength increases.

ASTM standard sieves can be obtained with wire diameters of lOOum and

above with respective spacings of lOOum and above. We used three

different sieves with the designation of #100, #70 and #50. The wire

dimensions and the spacing between the wires are presented in Table 1. The set-up is shown in Figure 1. Two slots were made in the bottom

rim of each of sieve such that a fiber may be laid parallel to one of

the sets of wires in the inverted sieve. This set-up is placed

between the launch and receiving optics of a spectral attenuation

measurement system. Care is taken to confirm the wire mesh is well

supported from underneath so that the mesh does not give way or flex

under weight of the loads. The fiber to be measured is installed in

the slots in the sieve and a reference measurement is made. The fiber

is then loaded with standard weights of 1 kg each. These weights are each a smooth 1" x 2" block of steel with a protective coating on one side (The protective coating prevents the fiber from being cut by the corners of the steel plate). Additional spectral scans are made and

compared to the reference scan to give microbend losses due to the

respective loads.

Results

Single mode optical fibers with varying deltas and cutoff wavelengths

were evaluated for microbending using this technique. The increase in

attenuation depends on the load on the fiber, type of mesh and

sensitivity of the fibers to microbending losses (i.e. how low the

delta and cutoff wavelength were)

.

Mesh #70 seems to give the most discriminating results with the single

mode fibers tested; bends are not too sharp and attenuations are not

overly high. The results presented in this report show effects of

Mesh #70 with loads upto 4 kg varying in steps of 1 kg load.

126
Mesh #100 on single mode fibers did not yield any increase in

attenuation. The mesh wires are extremely small and close together.

Apparently the effect of the force exerted by the wires is absorbed by

the coating. The effect is as if two smooth plates were squeezing the

fiber. Interestingly this does not change the attenuation of the

fiber. Further tests are planned to evaluate the effect due to Mesh

#100 and other meshes on bare/ uncoated fibers.

Further Experiments

Microbend analysis using this setup will also be performed on multimode fibers. These results will be presented at the conference. A

correlation between the current results and microbend analyses in

earlier technical papers will also be attempted.

References

1) Furuya and Suematsu (Applied Optics, Vol. 19, p. 1493,

1980)

2) J.Auge, P.Dupont, L.Jeunhomme (Bending and Microbending

Loss Sensitivity of Step Index Single Mode Fibers, 1984

Symposium on Optical Fiber Measuremnts, Boulder,

Colorado)

.

3) Proposed EIA Test procedure FOTP-68 (Microbend Test

Procedure)

.

TABLE 1

Mesh Wire Dimensions and Spacing

Mesh # Wire diameter Spacing

1) #100

2) #70

3) #50

110 urn 152 urn 215 urn

150 urn 212 urn 300 urn

127
LOADS

Figure . 1

SPECTRRL RTTENURTION

Cuto-f-f - 1180 nm nesh »70 Oe 1 t a = ? Loads 1 - 4 Kg

Wavelength (nms)

Figure. 3

128
PREDICTING MICROBENDING LOSSES IN SINGLE-MODE FIBERS

Neil Kamikawa

Naval Ocean Systems Center, P.O. Box 997, Kailua, HI 96734-0997

Ching-Ten Chang

San Diego State University and Naval Ocean Systems Center,

San Diego, CA

INTRODUCTION. Conventional techniques for evaluating microbending resistance in single-mode fibers using basketweave [1],

pin [2], drum [3], and sandpaper tests depend on excess loss

measurements. These tests can evaluate one fiber against

another, but the stresses induced in the fiber may not be representative of cable structures or other environments. An alternative evaluation method involves predicting excess losses using a model based on Petermann's microbending theory [4], which depends

on mode-field radius, and geometry of the microbends. This

method is demonstrated by predicting the excess losses in two

precision-wound spools of fiber.

MODEL. The model shown below for periodic, Gaussian-shaped

microbends that are separated by a distance L was derived from

Petermann's theory. It relates excess loss in dB/km to fiber

specifications and microbend geometry. Ye is the maximum amplitude and A is average half-width at the 1/e amplitude, as shown

in Figure 1. Variations in the half-width values can be described by a standard deviation, 0" . The electric field distribution of the fundamental mode is also approximated by a Gaussian

function with a radius W, which is defined at the 1/e value.

2 2 (B + (T )

129

13.644 Y„

LW

2 3 A B <T*B

2 2 S/2. (B + or )

+

2 2 3/2. 2(b + <r ) exp
where B = knW2//2~

k = 27T/X

n = refractive index of the core

RESULTS . To demonstrate the model excess losses were calculated and measured in two precision-wound spools of fiber described in Table 1. Precision winding of fibers produces crossovers that can be approximated by the Gaussian-shaped microbends.

Each layer of the precision-wound spool forms a helix with a lay

angle nearly 9 0 degrees with respect to the axis of the mandrel.

The next layer forms a helix in the opposite direction. As a result the fiber crosses over a fiber beneath it twice per turn

so that the period of the microbends are equal to half the circumference of the mandrel (L= "7C x 5.72 cm). The crossovers cause

perturbations on the fiber axis and coupling of the fundamental

mode to the lossy higher-order mode. This is assumed to be the

primary physical mechanism for the microbending loss.

Table 1. Fiber and winding specifications

Fiber 1 Fiber 2

LP,, cutoff X (jam) 1. 135 1.209

Mode-field radius (um)

@A = 1.3 um 4.87 4.73

@7v= 1.5 jam 5. 65 5. 32

@ A = 1.55 jam 5.96 5.55

Delta (%) 0.27 0. 30

Fiber OD (jam) 125 127

Coating OD (jam) 243 241

Length (km) 2.2 2 . 0

Mandrel diameter (cm) 11.43 11.43

Crossover amplitude (um) 35 35

Winding tension (grams) 200 200

Average crossover

half-width, A (mm) 1.05 1. 05

Standard deviation Q~ (mm) 0.201 0.210

The fibers were wound onto 11.43-cm-OD mandrels in 17 layers

under 200 grams of tension. Figures 2 and 3 illustrate the

measured losses in the fibers before and after winding. Fiber 2

130
exhibits better microbending resistance at long wavelengths due

to tighter mode confinement.

Y0 was calculated to be 35 urn based on the fiber OD, and A

was estimated to be 1.05 mm for both fiber spools by visual

observation. The standard deviation values, Q~ , are 2 01 urn and

210 urn for fibers 1 and 2, respectively. These values were

chosen for the best match between calculated and measured excess

losses at 1.3, 1.5 and 1.55 urn. Table 2 compares the excess

losses calculated using the model and measured excess losses at

the three wavelengths. Errors in the calculated values are

tentatively attributed to measurement errors in the mode-field

radius, and visual observation of A. Winding imperfections also

contributed to measurement and prediction errors.

Table 2 . Calculated and measured excess losses in dB/km

(Measurement errors are in parentheses.)

1.3 urn 1.5 urn 1.55 urn Fiber calc/measured calc/measured calc/measured

1 .017/. 001 (±.013) .090/. 086 (±.015) .220/. 202 (±.023)

2 .031/. 008 (±.015) .067/. 044 (±.022) .110/. 110 (±.031)

CONCLUSION. Prediction of excess losses as a method to

evaluate microbending resistance of single-mode fibers is a viable alternative to the other testing methods. Efforts are

underway to improve the predictive capabilities of the model by

including mechanical bending properties of the fiber.

REFERENCES

.

1. A. Tomita, P.F. Glodis, D. Kalish, P. Kaiser, Technical

Digest - Symposium on Optical Fiber Measurements, Boulder,

CO, p 89-92, 1982.

2. V.A. Bhagavatula, I00C '81, paper TUK5, San Francisco, CA,

27-29 April 1981.

3. D. Gloge, Bell System Technical Journal, p 245-262, Feb, 1975.

4. K. Petermann, Opt. and Quant. Elect., p 167-175, Sept, 1977.

131
E

0.2 0.0 -|—.—i—.—i—I—.—i—.—i—I—i—i—i—.—I—i—.—.—i—I—i—i—i—i—I—i—i—i—i—

1.10 1.20 1.30 1.40 1.50 1.60 1.70 WAVELENGTH (microns)

Figure 2. Spectral attenuation for fiber 1

132
ON THE NUMERICAL APERTURE OF A SINGLE-MODE FIBER

D. N. Christodoulides

L. A. Reith

M. A. Saifi

Bell Communications Research

Murray Hill, New Jersey 07974

Light-emitting diodes (LEDs) coupled to single-mode fibers (SMFs) may become the

first generation lightwave system to be deployed in the local distribution network.1 The

concept of the fiber numerical aperture (NA) which originates from geometrical optics is

commonly employed to estimate the coupling efficiency of an LED to a multimode fiber.

For a step-index fiber, NA = (n* -nf )

1/2

, where n± and n2 are respectively the refractive

indices of the fiber core and cladding. However, the problem of an LED coupled to a SMF falls within the frame of wave optics and therefore the concept of SMF NA must be

reconsidered. In this paper the NA of a SMF is revised so as to incorporate the wave

nature of the coupling process. This is done by interpreting the results of electromagnetic

theory from a geometrical point of view.

We envision the LED light emitting surface as an ensemble of classical radiators,

radiating in a mutually incoherent fashion. The power collecting ability of a SMF is then

investigated by embedding one of these radiators in an infinite SMF. Let this radiator be

a simple current dipole located at a distance r from the fiber axis as shown in Fig. 1.

Under steady-state propagation, the dipole supplies power to the fundamental mode LPqi,

in both the forward P0 + and backward direction P0 .. In order to convert the wave picture

to a geometrical one, it is assumed that the elementary dipole provides the power %T to the

LP01 mode only through a "cone" of angle he whose axis is parallel to that of the fiber.

The electric field within the SMF is described using the Gaussian approximation , ie.,

E = exp(-r2 /2ro ), where r0

is the mode field radius. It can then be shown that3

133
Or = exp(-ra /2rg) (1)

When the dipole is located outside the fiber, the maximum "acceptance angle" can be

obtained from Snell's law, ie., = (Xo/tr

r

0) exp(-rJ /2r£). We identify as the

numerical aperture NA of a SMF

NA = exp

r

2 *

it r0 exp

2r0

2 (2)

For step- index SMFs, the quantity Xq/vtq is very close to the geometrical optics NA.

Unlike multimode fibers whose NA depends on the index distribution, in the case of a

SMF, NA is only a function of wave parameters such as wavelength \j and mode field

radius r0 and also a function of radial position r = (at 2 +y

2 )

1' 2

.

Using (2), the coupling efficiency of an LED butt-coupled to a SMF can be obtained

from

JfdxdyJ J B(x,y;<fr, 6) sin 6 J8 d$

•n = T

0 0 2k nil (3)

// dxdy f f

B(x,y; <}>, 6) sin 6 dBd<&

o o

where T = 4n/(l+n)2 is the Fresnel transmission coefficient and B is the radiance of the

LED.

Butt-coupling efficiencies were measured at \) =1.3 jam using an edge-emitting LED

and a surface emitting LED. The fiber used was a step-index fiber whose mode field

radius at \j 1.3 ^ra was 3.5jim.

The measured butt-coupling efficiency for the edge-emitting LED was -19.4 dB, which

compared to a calculated value of -18.6 dB using (3). If the center of the edge-emitting

134
LED is located at (a,p) in the Cartesian system whose origin coincides with the fiber axis,

then equation (3) can be used to predict the coupling sensitivity due to misalignment. The

Y axis is perpendicular to the junction plane of the edge-emitter. Figure 2 shows the

coupling sensitivity S, and S, as a function of displacement a and 0 respectively. The solid

dots represent the experimental results. Note that the S, curve is a result of the wave

nature of the numerical aperture.

The butt-coupling efficiency for the surface -emitting LED was found to be -35 dB as

compared to -34.8 dB evaluated using (3).

REFERENCES

[1] J. L. Gimlet, M. Stern, R. S. Vodhanel, N. K. Cheung, G. K. Chang,

H. P. Leblanc, P. W. Shumate and A. Suzuki, "Transmission experiments at 560

Mbit/s and 140 Mbit/s using single-mode fiber and 1300 nm, LEDs", Electron.

Lett., vol. 21 , pp. 1198-1200, 1985.

[2] D. Marcuse, "Gaussian approximation of the fundamental modes of graded-index

fibers", J. Opt. Soc. Am., vol. 68, pp. 103-109, 1978.

[3] D. N. Christodoulides, L. A. Reith and M. A. Saifi, "Theory of LED coupling to

single-mode fibers", to be published.

135
Br. CLADDING

1 \-~

1 1 —I—

r

i I r^=C> Pot

JL.

Fig. 1. A current dipole within a single-mode fiber.

CORE

r»i

X

Fig. 2. Coupling sensitivities 5, and S, as a function of displacements a and p

respectively. The solid curves represent their behavior as predicted by eq. (3) and

the solid dots the experimental results.

136
OPTICAL FIBRE FAR FIELD ESI DETERMINATION USING THE EQUALISATION WAVELENGTH

A C Roucouval as , S C Robertson and J Walker

GEC Research Limited Hirst Research Centre East Lane Wembley Middlesex HA9

7PP United Ki ngdom

ABSTRACT

It is shown that measurement of the equalisation wavelength, (x e ) instead of

the LP11 mode cut-off wavelength (x c ), can be used to calculate a more reproducible and reliable value for the equivalent step index profile

parameter. Unlike X c , X e is virtually independent of fibre bending.

INTROD U CTION

When a single mode optical fibre is operated in the dual mode regime,

with an off-axis launch, power is guided in both the LPqi and IP\\ modes.

Provided the group time delay difference, T, between the modes is less than

the coherence time of the light t, (where t < X_2_ , 6X is the line width, c c6X~

the speed of light in a vacuum, and X the wavelength) then modal

interference will be observed [1,2].

The equalisation wavelength, X e

, of the dual mode fibre is the wavelength

at which the group velocities of the two interfering modes are the same. When observing the fibre in the near field, it is possible to measure x e of

the fibre precisely [3], by varying wavelength, and recording the intensity

oscillation of the two IP\\ lobes. This effect has

already been used [4,5] in fibre optic measurements to relate x c

, the cutoff

wavelength, to X e

, and further to determine the V-value of fibres. In this

work we extend this technique to the determination of reliable ESI fibre

parameters using X e

, as an alternative to the current practice of using x c for ESI measurements [6,9].

FAR FIELD METHOD

We have examined the possiblity of using the far field intensity

distribution measurement data in conjunction with X e as an alternative

method to the cutoff wavelength, X c

, for determination of the fibre ESI

parameters.

The method utilises the far field radiation pattern to determine the ESI core radius.

Usi ng

U

= (1+/2)V

l + ( 4+V4)l/4

or Newmann's approximation with V=VeX e /x, (V e

= 3.04) together with the far

field intensity distribution:

2 = U

2W

2

{

J

0

(a ) - aJi(a) VU )

}| (U 2

^t 2 )(W2

4xx 2

) UJKU)

I*

the only unknown ESI core radius can be found by a best fit of the above

formula to the measured data. Hence the ESI parameters can be calculated.

A typical best fit to the far field data using the above formula is shown

in Figure 1. The fibre has matched cladding and a small axial dip. The far

137
field pattern was measured for a number of wavelengths and the ESIs obtained were shown in Table 1 showing the i nsensi ti vi ty of the ESI measurments to

the wavelength used.

We can now write Petermann's spot size [9] for a step index fibre in

terms of X e as follows:

3/2 6

1 ^ = 0.634 + 0.30545 (X_) + 0.003648 (x_) - 0.0006506 (X ) 3 X g X ^ X g This allow us to determine the dispersion of the fibre, (knowing a, and

x e ). Figure 2 shows the measured and computed dispersion results for this

fibre type. The agreement is good considering that the fibre is quasi step

index. The zero material dispersion wavelength is predicted to within 5 nrn.

I nsensi t i vi ty o f x p to bends

It is well known that X c is dependent on the state of the fibre, i.e.

bending [5], affects the cutoff wavelength significantly. Since bends are normally unavoidable, we have decided to examine the dependence of X

e to

bending. For this, one turn of fibre was wound on cylinders of different

di ameters

.

Light from a Raman laser was launched into the fibre, after it has been

filtered by a monochromator, and the emerging light at the output (near

field) was displayed on a TV monitor. The oscillation of the LP\i mode

pattern was observed as the wavelength of the monochromator was scanned. The

wavelength positions for minima of one of the lobes was noted. Figure 3 shows the results plotted for a variety of cylinder diameters. We can see that as the bend radius decreases, the visibility of the oscillations is

gradually lost for long wavelengths. We have found in this case that

X e

= 1.064 urn and this remains constant to within 1 nm for bends down to

3.5 cm diameter. Below this measurement of X e is not possible. The results

show, that compared to X

c the measured value of X e is quite insensitive to

bends

.

Profile dependence of Ve/Vc

It was considered necessary to determine how the ratio of X c

/X

e is

dependent on the core refractive index profile. This was determined

computationally using graded index core profile (a value profile). Figure 4 shows the results of the ratio Ve /V c versus a, the profile grading power.

It is clear that the ratio is insensitive to the a value so long as a>5. In the case of a=5, the change in the ratio from that of a=°, (step index) is

10%. Since the manufactured fibres have a values >>4, (non-dispersion

shifted), the change of the ratio improves drastically.

This indicates that the measurement of X e is sufficient for the accurate

prediction of the theoretical X c even for graded index profiles.

CONCLUSIONS

A technique of determining the ESI parameters based on the measurement of

the equalisation wavelength of the LPn and LPqi modes has been developed.

The fibre dispersion is shown to be in good agreement with the measured

results for quasi step index core fibres. The technique has the advantage

that the measurement of X e is more accurate than the measurement of x c (which can be predicted from X e for a values > 4).

This method is therefore more reliable than determining the ESI

parameters by measurement of the cutoff wavelength of the LPn modes.

138
ACKNOWLEDGEMENT

Part of this work was supported by DTI under JOERS.

REFERENCES

1 PLANAS , S.A., BOCHOVE, E. , and SRIVASTA, R. : 'Geometrical

characterisation of liquid core fibres by measurement of thermally

induced mode cutoffs and interferece ' , Applied Optics, August 1982,

Vol 21, No 15, pp 2708-2715.

2 SHIBATA, N. , TATEDA, M. , SEIKAI, S., and UCHIDA, N. : 'Spatial

technique for measuring modal delay differences in a dual -mode

optical fibre', Applied Optics, May 1980, Vol 19, No 9, ppM1489-1492

3 McMILLAN, J.L.: 'Novel techniques for measurement of wavelength

of zero relative delay between modes of a dual -mode optical fibre',

Electronics Letters, 1983, Vol 19, pp 240-42

4 McMILLAN, J.L. and ROBERTSON, S.C.: 'Cutoff wavelength determination

in single mode optical fibres by measurement of equalisation

wavelength', Electronics Letters, 1984, Vol 20, pp 689-699

5 MCMILLAN J.L. and ROBERTSON, S.C.: 'Characterisation of single-mode

optical fibres by measurement of equalisation wavelength', Electronics

Letters, 1985, Vol 21, pp 295-296

6 MARCUSE, D.: 'Loss analysis of single mode fibre splices', Bell

Systems Tech., 1977, Vol 56, p 703

7 MILLAR, C.A.: 'Direct method of determining equivalent-step-index

profiles for monomode fibres', Electronics Letters, 1981, Vol 17,

pp 458-460

8 ANDERSON, W. : 'Consistency of measurement methods for the mode field

radius in a single mode fibre', IEEE Journal of Lightwave Technology,

April 1984, LT-2, pp 191-197

9 HUSSEY, CD. and MARTINEZ, F: 'Approximate analytic forms for the

propagation characteristics of single-mode optical fibres', Electronics

Letters, November 1985, 21, pp 1103-1104

FIGURE 1 TABLE 1

Fibre: GECMC

-t o :j 1, |

( " j ] ^ J "4" f-y

Wavel ength

(nm)

1312

1409

1555

1700

4.15

4.09

4.05

4.03

3ESI

(lim)

0.1256

0.1274

0.1287

0.1293

NA E si

139
G . E . C '. m . c .

12 13 I' '5 IS Wavelength ( )

Fiqure 2

_j 1 ' ' i i i 1.0 1.02 |.M KM 1.08 1.1 1.12 Wavelenqth (ym)

Fiqure 3

Alpha"v"Ve>Vc

»3.

1B

33

e|__ , , , , , , , V.2* lit O* 132 o« Uf Uf 1.11 1.11

Fiqure 4

140
Impact of Reflection on the Performance of Single-Mode Fiber Optic Systems

Shuenn Jyi Wang

Bell Communications Research, Morristown, NJ 07960

1. Introduction

The effect of reflection on a laser's characteristics has been discussed in many papers, but only a few articles

address the impact of reflection on the performance of fiber optic systems [1,2]. Reflecting surface and laser

facet form a cavity that acts as an external resonance cavity for the laser. The resonance cavity changes the

laser's lasing wavelength, Iinewidth, and threshold current, as well as enhances low and high frequency

noise[3,4]. All these effects result in output power fluctuation or undulation, mode hopping, and hence

degrade signal-to-noise ratio [5-7]. For a single-mode fiber optic system, large reflections can be introduced

by connectors that have an air-gap [1,6]. The maximum reflectivity on such connectors is 4r [6], where r is the

single-reflection coefficient between air and the fiber end. For a silica fiber with a refractive index of 1.46, r = 3.5%. Therefore, the reflectivity of connectors can vary from 0 to 14% depending on the length of the airgap. To examine the effect of reflection on fiber optic systems, we introduce different amounts of reflection

into the systems, and then measure the systems bit-error-ratio (BER). Twenty-two transmitters operating at

various bit rates and employing different kinds of laser diodes were checked. The result shows that the power

penalty due to reflection for single-mode fiber optic systems operating up to 570Mb/s could be as high as 1.9 dB for the lasers tested.

2. Experiment

Figure 1 illustrates the configuration of the fiber optic systems for the BER measurements. The transmitter

usually includes the multiplexers that multiplex several DS3 channels, each corresponding to 44.736 Mb/s, and some overhead bits. The multiplexed signal is scrambled and coded for a 50% duty cycle. The scrambled

signal then drives a laser diode that operates either in multi-longitudinal mode at 1.3 wm or in singlelongitudinal mode at 1.55 urn. A specially designed and calibrated reflection generator is used to control the amount of intended reflection [8]. The reflection generator is made with an attenuator and a segment of fiber

with dielectric coating on one end face (see Figure 1). It is then connected directly to the output of the

transmitter by a fiber jumper cable three meter in length. The adjustment of the reflection generator changes

the reflection amount as well as the forward attenuation. Hence, we use a low reflection attenuator to

compensate the adjustment of the reflection generator and control the received power for the BER

measurement. To minimize the discrepancy in the power reading due to connect and disconnect, we used a

coupler that was characterized before being connected to the power meter and the receiver. Also, the coupler

was held stationary until the measurement for a transmitter was finished. During the measurement period,

only the knobs of the attenuators were touched so that the polarization and the phase condition could be

141
preserved. For each transmitter measurement, some arbitrarily chosen DS3 channels are monitored. The

performance of these channels were not exactly the same. However, the difference of the BERs was not

significant and was consistent during the measurements.

The reflection amount was set at 1%, 6.3%, 15.9% and 22.4% (corresponding to 20, 12, 8, and 6.5 dB of

return loss, respectively) in the BER measurement for those transmitters with multi -longitudinal-mode lasers

operating at 1.3 urn. The 15.9% reflection is to simulate the case of an air-gap connector near the transmitter. We also took some transmitters' optical spectra corresponding to the above reflection setting to verify the

lasing wavelength's drift and the spectral width's change. For one transmitter, we inserted a 40 kilometer fiber

between the attenuator and the coupler to produce dispersion result. The measured BER then included the

effect of the mode-partition noise that is induced by reflection. For the transmitter with a distributed feed

back (DFB) laser working at 1.55 um region, we measured the BER versus the received power with the

reflection set at 1%, 6.3%, and 10%.

3. Results

Figure 2 shows the typical BER curves for a single-mode fiber optic system with reflections. Comparing the

data for the 15.9% and 1% reflection, we find about 0.05 to 1.9 dB power penalty for various transmitters. We did not control the polarization or the phase of the reflected field in the measurements. For those

transmitters that experience less power penalty, the phases or the polarizations of the reflected field and the

field in the laser cavity could be less matched. The leftmost curve in Figure 2 is the BER curve without

intended reflection, and is measured by removing the reflection generator. We find that, for some

transmitters, the system performance with 1% intended reflection is better than that without reflection. A

potential cause is that the polarization and/or the phase mismatch between the reflected field and the field in

laser cavity results in a weak optical feedback. The weak feedback then reduces some noise and narrows the

linewidth of each individual longitudinal mode. It was decided not to compare the data for the system with

and without intended reflection, because of the different physical configurations for these two measurements.

Although the root-mean-square (RMS) spectral width and the central wavelength for the systems using a

multi-longitudinal mode laser are changed by reflection, these changes are too trivial to affect the dispersion

power penalty. On the other hand, the mode-partition noise induced by reflection could be serious [5,9]. Using a long fiber, it is expected to see more dispersion power penalty caused by the induced mode-partition

noise. Figure 3 shows 0.45 dB power penalty for a 405 Mb/s single-mode fiber optic system with 15.9%

intended reflection and an additional 0.6 dB power penalty when using a 40 kilometer fiber.

All the measured data is summarized in Figure 4. The power penalty due to reflection is defined as the

difference between the received powers for a fiber optic system operating at BER = 10-8 with 1% and 15.9%

of reflection. For the system using a DFB laser, the data of 0% and 10% reflections is compared (see Figure

5).

142
4. Conclusion

It is found that reflection does affect the system performance for a single-mode fiber optic system. From the

data it is concluded that;

a. Different type laser diodes may respond to reflection differently [1]. However, for the same laser diode,

the power penalty due to reflection is determined primarily by the condition of "polarization" and

"phase" of the reflected field.

b. The central wavelength's drift and the spectral width's change due to reflection for a multi-longitudinalmode laser are insignificant. However, the mode-partition noise induced by reflection is not negligible.

c. For the fiber optic systems operating at higher bit rates, the degradation of the system performance due

to reflection becomes more serious.

Acknowledgment: The author would like to thank David S. Burpee and Nim K. Cheung for their stimulating

discussions, William C. Young for making and calibrating the reflection generator, Edward Cherkassky,

Garvey C. Ince, Sunil G. Joshi, and Dan L. Lucas for their help in collecting the data. REFERENCES

1. N. K. Cheung, SPIE Conference, Arlington, Virginia (1984).

2. V. J. Mazurczyk, Electron. Lett. 17, 143 (1981).

3. M. Fijiwara, K. Kubota and R. Lang, Appl. Phys. Lett. 38, 217 (1981).

4. K. Kikuchi and T. Okoshi, Electron. Lett. 18, 10 (1982).

5. N. A. Olsson, W. T. Tsang, H. Temkin, N. K. Dutta, and R. A. Logan, J. Lightwave Technol. LT-3,

215 (1985).

6. R. E. Wagner and C. R. Sandahl, Appl. Opt. 21, 1381 (1982).

7. G. A. Acket, D. Lenstra, A. J. denBoef and B. H. Verbeek, 1EEEJQE-20, 1163 (1984).

8. W. C. Young et al, to be published.

9. G. P. Agrawal and T. M. Shen, J. Lightwave Technol. LT-4, 58 (1986).

DS3 Transmitter Reflection Low-reflection Coupler Generator Attenuator

Receiver Error

Detector

}

-EJDielectric Coating

(Reflecting Point)

Optical Power Meter

Figure 1

143
J1

i

\ \\

1

1

1E-3

-44 -43 -42 -41 -40 -39 -38 -37 RECEIVED POWER CdBm)

Dg. 2 fl typical fiber optic system tilth B'/AX), U(0),

B.3X(+), 15.95a"), tnd Z2.4*(a) reflections.

-36

IE -4

IE -5

1E-6

IE-7

IE -8

IE -9

1E-10

IE- 1 1 1E-12

A = /.3a

-v^ a^

V

v

1 \ \

\ VVV

V

1

1 1 -42 -4 1 -40 -39 -3B -37 -36 -35 RECEIVED POWER CdBm)

-34

Fig. 3 R 405 Mb/s -Tiber optic system with U(0) tnd 152(0) reflections, tnd (X) using t 40 Km fiber (15. 9X reflec.

FIBER OPTIC SYSTEM BIT RATE (rib's) RECEIVED POWER CdBm)

l.

1 4 The pouer r rpenalty due to reflection for c , _

» f,g 5 1^,, fib tr optic system using DFB ltsir Kith single-mode fiber optic systems. Bx(x)i lxm> 6 . 3X ( +)i lnd 1BX(B ) r ,f|, c tlons.

144
Spurious Edge Elimination for Accurate Glass Geometry Measurements

Devon R. Mcintosh and David A. Hall

Corning Glass Works

One method of performing glass geometry measurements on optical waveguides is to

digitize the near-field image of the fiber end using a gray scale analyzer. By defining a

certain percentage level of the gray scale value as the transition threshold, fiber image

edge coordinates can be determined. These edge coordinates are the data points

used in the calculation algorithms suggested by CCITT for glass geometry

measurements using such a raster scan. Clad diameter is determined by the diameter

of the best-fit circle to the clad edges. Non-concentricity is determined as the distance

between best-fit circle centers of the clad and the core; and the radii from these

centers are used for noncircularity calculations.

In single mode fibers, due to the step-like changes of the index of refraction, any

percentage threshold level can theoretically be used. In practice, the 50% threshold

level has given the best results. In multimode fibers, with their associated graded

index profile, glass geometry parameters for the core must be defined close to the

baseline of the nearfield intensity patterns. At such a low threshold a significant

number of spurious edges are detected due to noise or foreign matter causing an unwanted threshold crossing. These edges must be identified and discarded before

applying the calculation algorithmns in order to maintain acceptable accuracy and

repeatability.

A method of identifying these spurious edges has been developed which compares

edge locations to a mathematical model of the fiber end and its interaction with the

measurement system. This model is based on the following assumptions:

1 . The number of edges belonging to the fiber image is more than the

number of spurious edges.

2. The image's deviation from circular is either elliptical or oval or some

combination of both. (Although the mathematics is derived using this

assumption, the results can be used to eliminate spurious edges from

other somewhat circular shapes.)

3. This deviation can be modeled by some combination of sinusoids having

a period of % and 2k.

4. The location of the center of the best-fit circle at the 50% threshold level is a close approximation to the center at the measured level.

Using the above assumptions, all edges belonging to the image of the multimode core can be analyzed for their conformity to the mathematical model. Measurement system

error estimates are used to determine the tolerance around the model parameters.

Since these parameters can be calculated for each edge; any edge parameter falling

outside the tolerance can be discarded.

145
The mathematical model for determining edge parameters is derived as follows:

Given a data distribution for the core edges as shown in fig. 1 and using the

center as stated in assumption 4, the distribution of radii is mapped and the

best determined radius, is determined as shown in fig. 2. For each data

point:

Let: f(0) = R(0) - (Local deviation from the circle of radius R^) (1)

and: f '(6) = (R(9) - R(6')) / (0-0') (Local slope of the data) (2)

where R(8) and R(6') are adjacent radii and (0-0') is the angle

separating them in radians.

For purely oval deviations (modeled as a sinusoid with period 2k and

maximum amplitude A-j

:

f(0) = A-, sin 0

and f '(©) = df(0)/d0 = A1 cos 0. Since (A-, sin 0)

2 + (A-| cos 0)2 = A-j 2

then (f(0)) 2 + (f'(0))2 = A-| 2

. (3)

For purely elliptical deviations (sinusoid of period 2k and max. amplitude A2 )

f(0) = A2sin 20

and f *(0) = df(0)/d0 = 2A2 cos 20.

As above:

((A2 cos 20) / 2)

2 + (A2

sin 20) 2 = A2

2

and (f '(0) / 2)

2 + (f(0)) 2 = A22

. (4)

Since non-circularity deviations will be a linear combination of ellipticity (of

maximum magnitude a2) and ovality (of maximum magnitude a-j) separated

by a phase angle (cp);

f(0) = a-| sin 0 + a2

sin (2(0+9)) (5)

and f '(0) = a-| cos 0 + 2a2 cos (2(0+cp)) (6)

Eqn. (3) then becomes:

(f(0)) 2 + (f(0))

2 = A1 2

(0) (7)

and Eqn. (4) becomes:

(f'(0)/2) 2 + (f(0)) 2 = A2

2

(0) (8)

146
It is important to note that there are now two different definitions for f(0) and f '(9)- Eqn's

(1) and (2) give the physically measured definition and Eqn's (5) and (6) give the

mathematically modeled definitions. This in turn gives two different definitions for

A-j 2

(0) and A2

2

(0). To avoid confusion the subscript m will be used for the

mathematically modeled definitions and the subscript p will be used for the physically

measured definitions.

An average of A-|

p

2

(0) and A2p

2

(0) is needed with which to compare each individual

parameter. Therefore let A-jp2 and A2p

2 be the averages of A-|

p

2

(0) and A2p

2

(0)

respectively.

Spurious edges may now be eliminated by determining the tolerance around A^2

and A2p

2 through the propagation of system error estimates. A-|

p

2

(0) and A2p

2

(0)

must be within this tolerance for an associated edge coordinate to be retained. a-\ and

a2

in Eqn's (5) and (6) are needed for this calculation and are determined as follows:

Let: A 1m

2 =

1 /27l jQ2* A1 m2

(0) d0 (Average of A1 m

2

(0)) (9)

and A2m2 = 1 /27J0

2* A2m2 (0)d0 (Average of A2m

2

(0)) (10)

After substituting Eqn's (5) and (6) into Eqn's (7) and (8), and then into Eqn's (9) and

(10), and then integrating; the following results are obtained:

3,2 = 4.44 (A2m2-AA1m2 )

Given a correspondence between A2m

2 and A2p2

, and between A-j m

2 and A-jp2

, the following is also true: ai

2 = 4.44 (A2p2 -AA 1p

2

) a2

2 = 1.11 (A 1p2 ^.6A2p2

)

The expected tolerance around Afp2 and A2p

2 comes primarily from three sources;

edge detection errors, the estimation of Rp (Fig. 1 and Fig. 2), and the functional

dependence of A-j p(0) and A2 p(0) on 0. The first two sources are measurement errors and can be estimated. The third source can be analytically determined from Eqn's (9)

and (10). All three are propagated through the mathematical model (defined by a-| and

a2) to determine tolerances for A

-jp

2 and A2p

2.

147
Let: T^-j = Tolerance around A-jp^

TA2 = Tolerance around A2p

2

Every retained edge has to satisfy the following inequalities:

A 1p

" TA1 * A1p(0) * A 1p + TA1 , r

'

A2p

" TA2 * A2p(0) * A2p +TA2"

Glass geometry parameters are then computed for the cleaned image using the

standard algorithms. Repeatabilities on the same order as for single mode fiber have

been obtained on multimode fiber using this technique.

Fig. 1 Collection of core edges at a low threshold level

with best determined radius (R^).

b Radius

Fig. 2 Distribution of radii with best determined radius (R^)

148
INDEX OF AUTHORS

W. T. Anderson 1,49,65

R. C. Alferness 115

R. D. Birch 107

L. Bosselaar 15

A. C. Boucouvalas 137

M. C. Brain 73

E. L. Buckland 31

L. Button 11

M. Calzavara 53

C. -T. Chang 129

D. N. Christodoulides.... 133

G. Coppa 35,53

E. A. Cottrell 73

R. Crotts.... 81

L. Curtis 57,69

H. Damsgaard 23

P. Di Vita 35,53

C. K. Eoll 125

R. Engel 43

M. E. Fermann 77

S. Furukawa 89

A. 0. Garg 125

D. A. Hall 145

0. Hansen 23

W. H. Hatton 31

R. M. Hawk 11

D. R. Hjelme 85

T. Ito 101

Y. Iwamoto 19

W. S. Jackman 93

A. J. Johnson 65

Neil Kamikawa 129

F. Kapron 81,97

J. P. Kilmer 49,65

Y. Koyamada 89

C. Kozikowski 81

G. Kuijt 15

K. Kurumada 101

W. F. Love 11

R. S. Lowe 121

P. D. Ludington 61

F. Martinez 77

P. Matthijsse 15

D. R. Mcintosh 145

E. McNair 43

A. R. Mickelson 85

F. Mogensen 23

Y. Namihira 19

M. Ni shimura 31

D. N. Payne 77,107

D. L. Philen 7,27

S. B. Poole 77,107

M. Potenza 35

Catherine M. Ragdale 97

M. R. Ramadas 39

W. A. Reed 7 L. A. Reith 133

S. C. Robertson 137

M. A. Saifi 133

C. Saravanos 121

D. Schickentanz 93

V. Shah 57,69

M. R. Shenoy 39

M. Tateda 89

E. A. Thomas 65

K. Thyagarajan 39

J. F. van Luijk 15

M. P. Varnham 107

I. P. Vayshenker 85

J. Walker 137

Shuenn Jyi Wang 141

L. Wei 121

Ian A. White 61

Katsuya Yamashita 89

W. C. Young 57,69

149
NBS-114A irev. 2-8Q

U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

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1. PUBLICATION OR

REPORT NO. NBS/SP-720

2. Performing Organ. Report No. 3. Publ ication Date September 1986

4. TITLE AND SUBTITLE

Technical Digest, Symposium on Optical Fiber Measurements, 1986

5. AUTHOR(S)

G. W. Day and D. L. Franzen, Editors

6. PERFORMING ORGANIZATION (If joint or other than N8S. see instructions)

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

7. Contract/Grant No.

8. Type of Report & Period Covered

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City. State. ZIP)

Sponsored by the National Bureau of Standards, in cooperation with the

IEEE Optical Communications Committee and the Optical Society of America

10. SUPPLEMENTARY NOTES

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11. ABSTRACT (A 200-word or less (actual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here)

This digest contains summaries of 34 papers presented at the Symposium on Optical Fiber Measurements, held September 9-10, 1986, at the National Bureau of

Standards, Boulder, Colorado.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolons)

fiber optics; instrumentation; measurements; optical fiber; reviews

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