A1110E TD3bMM

NBS

PUBLICATIONS

NATL INST OF STANDARDS & TECH R.I.C.

A1 11 02903644

Symposium on Optical/Technical digest - QC100 .U57 N0.748 1988 V1988 C.2 NBS-PUB ^lAL PUBLICATION 748

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

Technical Digest

Symposium on Optical Fiber

Measurements, 1988

Sponsored by the National Bureau of Standards

in cooperation with the IEEE Optical Communications Committee

and the Optical Society of America
m he National Bureau of Standards 1 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 to assure international competitiveness and leadership of U.S.

industry, science arid technology. NBS work involves development and transfer of measurements, standards and related

science and technology, in support of continually improving U.S. productivity, product quality and reliability, innovation

and underlying science and engineering. The Bureau's technical work is performed by the National Measurement

Laboratory, the National Engineering Laboratory, the Institute for Computer Sciences and Technology, and the Institute for Materials Science and Engineering.

The National Measurement Laboratory

Provides the national system of physical and chemical measurement;

coordinates the system with measurement systems of other nations and

furnishes essential services leading to accurate and uniform physical and

chemical measurement throughout the Nation's scientific community,

industry, and commerce; provides advisory and research services to other

Government agencies; conducts physical and chemical research; develops,

produces, and distributes Standard Reference Materials; provides

calibration services; and manages the National Standard Reference Data

System. The Laboratory consists of the following centers:

• Basic Standards2

• Radiation Research

• Chemical Physics

• Analytical Chemistry

The National Engineering Laboratory

Provides technology and technical services to the public and private sectors

to address national needs and to solve national problems; conducts research

in engineering and applied science in support of these efforts; builds and

maintains competence in the necessary disciplines required to carry out this research and technical service; develops engineering data and measurement

capabilities; provides engineering measurement traceability services;

develops test methods and proposes engineering standards and code

changes; develops and proposes new engineering practices; and develops

and improves mechanisms to transfer results of its research to the ultimate

user. The Laboratory consists of the following centers:

Applied Mathematics

Electronics and Electrical

Engineering2 Manufacturing Engineering

Building Technology

Fire Research

Chemical Engineering3

The Institute for Computer Sciences and Technology

Conducts research and provides scientific and technical services to aid

Federal agencies in the selection, acquisition, application, and use of

computer technology to improve effectiveness and economy in Government

operations in accordance with Public Law 89-306 (40 U.S.C. 759),

relevant Executive Orders, and other directives; carries out this mission by

managing the Federal Information Processing Standards Program,

developing Federal ADP standards guidelines, and managing Federal

participation in ADP voluntary standardization activities; provides scientific and technological advisory services and assistance to Federal agencies; and

provides the technical foundation for computer-related policies of the

Federal Government. The Institute consists of the following divisions:

Information Systems Engineering

Systems and Software

Technology

Computer Security

Systems and Network

Architecture

Advanced Computer Systems

The Institute for Materials Science and Engineering

Conducts research and provides measurements, data, standards, reference

materials, quantitative understanding and other technical information

fundamental to the processing, structure, properties and performance of

materials; addresses the scientific basis for new advanced materials

technologies; plans research around cross-cutting scientific themes such as nondestructive evaluation and phase diagram development; oversees Bureau-wide technical programs in nuclear reactor radiation research and

nondestructive evaluation; and broadly disseminates generic technical

information resulting from its programs. The Institute consists of the

following Divisions:

• Ceramics

• Fracture and Deformation3

• Polymers

• Metallurgy

• Reactor Radiation

"Headquarters and Laboratories at Gaithersburg, MD, unless otherwise noted; mailing address

Gaithersburg, MD 20899.

JSome divisions within the center are located at Boulder, CO 80303.

'Located at Boulder, CO, with some elements at Gaithersburg, MD
Technical Digest—Symposium on Optical

Fiber Measurements, 1988

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 20-21, 1988

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, C. William Verity, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

Stimulating America's Progress 1913-1988

Issued September 1988
Library of Congress Catalog Card Number: 88-600561

National Bureau of Standards Special Publication 748

Natl. Bur. Stand. (U.S.), Spec. Publ. 748, 202 pages (Sept. 1988)

CODEN: XNBSAV

U.S. GOVERNMENT PRINTING OFFICE

Washington: 1988

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

It is our pleasure to present here summaries of all of the papers given

at the fifth biennial Symposium on Optical Fiber Measurements. Forty-two

papers, eight invited and thirty-four contributed, are included. This is a substantial increase over the 1986 Symposium, made possible by the elimination

of workshops from the program.

As always, certain themes appear in the program. This year, we see the

largest ever group of papers on optical time domain reflectometry (OTDR)

,

apparently motivated in large part by concern about reflections and their

effects on source behavior. The second largest group of papers concerns the

measurement of cut-off wavelength. And for the first time, we see a

significant number of papers on the characterization of planar optical

waveguides

.

The Symposium continues to enjoy broad international support. This

year's program includes participants from 25 organizations in 8 countries;

more than 40 percent of the papers are from outside the United States.

G. W. Day

D. L. Franz en

Boulder, Colorado

September 1988

iii
SYMPOSIUM COMMITTEE

L. Franzen, NBS , General Chairman

. W. Day, NBS, Program Chairman

H. Cherin, AT&T Bell Laboratories

P. DiVita, CSELT

R. L. Gallawa, NBS

A. H. Hartog, York Technology

J. T. Kane, Corning Glass Works

. P. Kapron, Bell Comm. Research

P. Lovely, Photon Kinetics

M. Ogai, Furukawa Electric

A. Reed, AT&T Bell Laboratories

P. R. Reitz, AMP

iv
CONTENTS

Page

Preface iii

Conference Committee iv

Advances in OTDR (Invited)

Mitsuhiro Tateda and Tsuneo Horiguchi, NTT 1

Investigation of Optical Components in Micrometer Range using

an OTDR System with the Balanced Heterodyne Detection

R.P. Novak, H.H. Gilgen, and R.P. Salathe, Swiss PTT

W. Hodel, P. Beaud, J. Schutz, and H.P. Weber,

University of Berne 7

155 km Long-Range OTDR Measurement at 1.55 /am Wavelength Region

Y. Namihira, Y. Horiuchi, and H. Wakabayashi , KDD

T. Oshimi, K. Kitagawa, and T. Ooka, Anritsu 11

Bidirectional OTDR Measurements Utilizing an Improved Folded-Path

Technique

E.L. Buckland and M. Nishimura, Sumitomo Electric 15

OTDR Experiment for Glass Optical Waveguides by Using

Interferometric Technique with 1.3 //m SLD

K. Takada, H. Nagai , N. Takato, and J. Noda, NTT . 19

Reflectance Measurement in Lightwave Systems: A Comparison of

Various Techniques (Invited)

S.K. Das, A.F. Judy, G.M. Alameel, R.M. Jopson, and T.F. Adda,

AT&T Bell Laboratories 25

Calibration and Use of an Optical Time Domain Reflectometer

for Measurements of Reflectance (Return Loss) of Optical

Fiber Components

Paul R. Reitz, AMP 31

OTDR Measurements of Optical Return Loss

Felix P. Kapron, Edward A. Thomas, and John W. Peters,

Bellcore 35

Return Loss Studies of a Single-Ended Single-Mode Fiber: Effect

of the Nonflat Fiber Endface

V. Shah, W.C. Young, and L. Curtis, Bellcore 39

Optical Fiber Up -Taper Backreflector for Laser Transmitter

Evaluation

Alan F. Benner, Daniel A. Fishman, and Herman M. Presby,

AT&T Bell Laboratories 43

v
Measurements of Polarization Mode Dispersion in Optical Fibers

(Invited)

Craig D. Poole, AT&T Bell Laboratories 47

Heterodyne Fiber Polarimetry with Stokes Parameter Detection

R. Calvani, R. Caponi , and F. Cisternino, CSELT . 53

Complete Birefringence Characterisation of Lithium Niobate

Integrated Optic Polarisation Controllers

G.R. Walker and N.G. Walker, British Telecom Research

Laboratories 57

Techniques for Refractive Index Measurements on Semiconductor

Waveguides: Prism Coupling and Modal Cut-Off Spectroscopy

C. DeBernardi, M. Meliga, S. Morasca, and B. Sordo, CSELT ... 61

Electrically Calibrated Photothermal Deflection Measurement

for Separating Channel Waveguide Loss Mechanisms

Robert K. Hickernell, Donald R. Larson, and

Robert J. Phelan, Jr., NBS 65

Fiber Measurements for Long Term Reliability: Strength,

Hydrogen, and Radiation (Invited)

F.V. DiMarcello, P.J. Lemaire , and J.R. Simpson,

AT&T Bell Laboratories 69

Fluoride Optical Fiber Measurements: Techniques and Status

(Invited)

Lynda E. Busse, Naval Research Laboratory 79

Bend Loss Measurements on Single Mode Fibers

A. Hordvik, M. Eriksrud, and S. Lauritzen, Optoplan 85

Induced Losses in Strained Loose Tube Fiber Cables

L. Bjerkan, E. Nesset, and 0. Aim, EB Cables. 89

Measurements of Single Mode Fiber Attenuation over an

Extended Spectral Range

Z. Pasturczyk, C. Saravanos , and R.S. Lowe, Northern

Telecom Canada, Ltd 93

Mode Transition Matrix Techniques for Multimode Fiber -Optic

Components (Invited)

Udo B. Unrau, Technischen Universitat Braunschweig 97

A Single Launch Technique for Determination of Mode

Transfer Matrices

S. Yang, D.R. Hjelme, I. P. Januar, I. P. Vayshenker, and

A.R. Mickelson, University of Colorado 103

vi
Sensitivity of the Mode Transition Matrices to Variations of

the Input Power Vectors

T.J. Gryk, G.E. Holmberg, and H.M. Blount, Naval Underwater

Systems Center : 107

Evolution of Power Distributions in Fiber Optic Systems - Development of a Measurement Strategy

M.J. Hackert, Corning Glass Works Ill

Accurate Measurement of Optical Power in Fibre Optic Systems

D.H. Nettleton, National Physical Laboratory 115

An Accurate Method of Determining Single -Mode Fiber Cutoff

Wavelength

C.N. Okafor and D.W. Peckham, AT&T Bell Laboratories 119

A Technique to Estimate the Cutoff Wavelength Profile in

Single Mode Fibers Using a Switchable Dual Wavelength OTDR

K.W. Kowaliuk and J. Ferner, Northern Telecom Canada, Ltd.. . . 123

Predicting the Cutoff Wavelength of Cabled Single Mode Fibers

from Cable Structure

W.H. Hatton and M. Nishimura, Sumitomo Electric 127

Simulated Cable Deployment for Cable Cutoff Wavelength

Measurements

J. P. Kilmer, R.M. Kanen, and W.T. Anderson,

Bell Communications Research 131

Cutoff Characteristics of Dispersion-Shifted and

Dispersion-Unshifted Fibers

V. Shah, L. Curtis, and W.C. Young, Bellcore 135

1300 nm Fiber Characterization by LP 01 and LP 02 Interferogram

Equalization Wavelength Measurement

FranCois Gonthier, Suzanne Lacroix, Richard J. Black,

and Jacques Bures, Ecole Po lytechnique , Montreal 139

Fiber Geometry, a Fiber Manufacturer's Viewpoint (Invited)

David A. Hall and Robert W. Smith, Corning Glass Works 143

Non-Contact, Interferometric Determination of the Outside

Diameter of Optical Fibers

M.J. Saunders, AT&T Bell Laboratories 149

On the Development of a Calibration Standard for Fibre

Geometry Measurement

J.G.N. Baines and K.W. Raine , National Physical Laboratory

A.G. Hal lam, York Technology 153

vii
An Electronic Industries Association Interlaboratory

Comparison to Resolve Differences in Multimode Fiber

Numerical Aperture Measurements

A.H. Cher in and E.D. Head, AT&T Bell Laboratories

D.L. Franzen and M. Young, NBS

M. Hackert, Corning Glass Works 157

Novel Interferometric Dispersion Measurement System

A.O. Garg, M.A. Brandtner, and C.K. Eoll, Siecor Corp.. .... 161

Refractive Index Profiling State of the Art (Invited)

K.W. Raine, J.G.N. Baines , and D.E. Put land,

National Physical Laboratory 165

Reconstructed Refractive Index Profile of Optical Fiber

by Refracted Rays

Nobuo Amano and Ken-ichi Noda, Tokyo University of

Agriculture and Technology 171

Calibration of Optical Fiber Preform Refractive Index Measurement

D.A. Svendsen, G. Beresford, and D. Marcus -Hanks

,

York Technology 175

Interlaboratory Comparison of Far-Field Methods for Determining

Mode Field Diameter Using both Gaussian and Petermann Definitions

Timothy J. Drapela, NBS 179

Improvements to the VAMFF System for Accurately Measuring the

Mode-Field Diameter of Dispersion Shifted Fibers

Jerry R. Parton, Corning Glass Works 183

Accuracy of Mode Field Diameter Measurements in Single-Mode

Fibres from Near-Field Technique

M. Artiglia, G. Coppa, P. Di Vita, and M. Potenza, CSELT. . . . 187

Index of Authors 191

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.

viii
Advances in OTDR

Mitsuhiro Tateda and Tsuneo Horiguchi

NTT Transmi sss i on Systems Laboratories

Tokai-mura, Ibaraki-ken, 319-11 Japan

1 . I ntroduct i on Barnoski and Jensen (1) first demonstrated

that detection and analysis of Rayieigh backseat tered light from

an optical probe pulse propagating down a fiber gives the

attenuation characteristics of the fiber. Using this technique,

we can also evaluate the splice losses and fault location in an

optical transmission line, composed of several optical fiber

cables, at one endface of it no nd es t r uc t i ve 1 y . Hence this

technique, so called optical time domain reflectmetry (OTDR), has

become a strong and helpful tool in all the fields of

manufacturing, testing and installation of optical fiber cables.

The gated detection technique greatly increased the ability of

the OTDR techn i que (

2

^ » O ^ , which enables the detection of weak

scattered light preventing detector sensitivity saturation by the

strong Fresnel reflection light.

2. Recent progress in the OTDR tech nique

( 1 ) En 1 argeme n t of the dynamic ra nge The OTDR technique has

enlarged its application target from multimode fibers at a short

wavelength (-0.8 p. m) to singlemode fibers at a longer wavelength

(1.3-1.5/i m) . However, the backseat tered light in singlemode

fibers is as small as a tenth of that in multimode fibers. To

overcome this difficulty, two kinds of approach have been

developed to enlarge the dynamic range. First, to increase the

input optical power (approach A) and second to decrease the

minimum detectable optical power (approach B) . They are listed

in Table 1 after Healey^. In approach A, various kinds of high

1
power laser light sources such as Q-switched Nd : YAG laser (5 ^(A.

= 1.32// m), Er + + + :glass 1 aser (6) (A =1.55// m) and P2O5 doped fiber

Raman laser (7) ( A =1.59// m) are used. In the Raman OTDR

technique (8) , Rayleigh backseat tered lights accompanied by Stokes

lights with wavelength ranging from 1.2 to 1.7 // m converted from

the strong input pumping light in the fiber under test is

detected (Fig. 1). Hence, this technique is used only for fault

location. Effectively improved optical signal power is utilized

in the correlation technique by signal processings^ . In

approach B, cryogenic detectors and data-averaging processing are

effectively used. A semiconductor laser is used both as an

optical pulse generator and as a backseat tered optical signal

amplifier, without a directional coupler, whose operation mode is

switched by a drive circuit (Fig. 2) (10 ^. In the photon counting

technique, photons in the b acks c a t t e r ed light are detected

digi tal ly^ 11 ^

( 12 * . In the heterodyne detection me thod ( 1

* "

1

^

,

we can expect to approach the shot-noise limit by increasing the

local oscillator power (Fig. 3d 3 )), although problems of fading

and polarization noise are h i gh- 1 i gh ted

( 2 Enhancement in res ol ution Small resolution (A 5mm)

significant in fault location is obtained by a narrow optical

pulse (t =46 ps). (16) However, 1/N times narrow pulse (t /N)

requires an N times broad bandwidth with an N times large noise

power, and it also diminishes the backseat t ered optical power to

1/N, which result in dynamic range decrease (7.5 log N dB).

Resolution as high as 15 11 m was reported by nonlinear detection,

in conjunction with a mode-locked dye 1 as er

1

F

£ • 4). An

optical ref 1 ec tome ter with a low coherence light source realizes

subm i 1 1 i me t er ( 1 8 ) and m i c r 0 me t e

( 1 9

}

resolution in the

application area of optical circuit diagnostics (Fig. 5).
(3) Reduction of single mode fi ber OTDR no ise Besides the weak

Rayleigh backseat tered light, single-mode fiber OTDR has the

problem of polarization dependent amplitude fluctuation due to

small birefringence in the fiber under test. An acous toopt i cal

deflector without polarization dependence^) anci polarization

s cr amb 1 er ^ 20 ) are effective in preventing this. Birefringent

launching fiber^l) composed of polarization maintaining fiber(s)

drastically diminishes the fluctuation (Fig. 6). The fading

phenomenon intrinsic to a coherent light source is reduced by a

nonlinear signal averaging operat i on ( 22 ^

(4) Us er frie ndliness As the potential of the OTDR technique

has become recognized, requests from users, for ease of

operation, have been increasing. OTDR for both 1.3 fj. m and 1.5

p, m fulfills one need in testing wideband optical fibers.

Automated OTDR waveform processing, eliminating cursor handling

procedure, is attractive because of constant measurement

repeatability by a nonskilled person, and the possibility of

remote controlled measurement (Fig. 7). The key point is in

judging the splice points, where time, and space averaging

processes, least square approximation and difference calculas

processes are used.

3 . Fu t

re t e chno 1

gy We have already realized the first stage

of the optical transmission systems of trunk lines(23) 4 j S

expected that the second stage, that is the era of the optical

subscriber lines, will begin in the near future, where several

demands for the OTDR technique will be emphasized. For example,

diminishing the dead zone will be essentially important in

diagnosing complicated optical networks such as Local Area

Networks. Furthermore, in testing the broadcasting type

branching networks, light source power for testing will be
limited so as not to destroy the adjacent equipment connected to

the system. To overcome these problems, new techniques in a

wider field must be developed for the OTDR technique.

References

(1) M. K. Barnoski et al . , Appl. Opt., 15, 2112, 1976.

(2) S. D. Personick, Bell Syst. Tech. J., 56, 355, 1977.

(3) T. Horiguchi et al . , J. Lightwave Tech. LT- 2 . 108, 1984.

(4) P. Healey, J. Lightwave Tech. LT-3 . 876, 1985.

(5) M. P. Gold et al . , 10th ECOC, 128, 1984.

(6) M. Nakazawa et al . , Optics Lett., 9, 312, 1984.

(7) K. Suzuki et al . , Electron. Lett., 2 2, 1273, 1986.

(8) K. Noguchi, J. Lightwave Tech., LT-2 . 1, 1984.

(9) J. J. Bernard et al ., 10th ECOC, 84, 1984.

(10) K. Suzuki, Electron. Lett., 2_0, 714, 1984.

(11) P. Healey, Electron. Lett., 17_, 751, 1981.

(12) B. F. Levine et al . , Electron. Lett., 21, 194, 1985.

(13) S. Wright et al . , 9th ECOC, 177, 1983.

(14) S. Wright et al . Optical Fiber Sensors Conference, 347, 1984.

(15) P. Healey et al .

, Electron. Lett., 18, 862, 1982.

(16) "A technical note on 00S-01 (in Japanese)", Hamamatsu

Photonics Inc., Hamamatsu, Japan, 1987.

(17) J. J. Fontaine et al . , Opt. Lett., 6, 405, 1981.

(18) K. Takada et al., Appl. Opt., .26, 1603, 1987.

(19) B. L. Danielson et al . , Appl. Opt., 26, 2836, 1987.

(20) K. Aoyama et al., IEEE, J. Quantum Electron., QE-1 7 , 862,

1981 .

(21) T. Horiguchi et al . , J. Opt. Soc. Amer. A, 2, 1698, 1985.

(22) P. Healey et al . , Electron. Lett., 20, 360, 1984.

(23) T. Uenoya, Japan Telecommunication Review, 101, 1985.

4
Table 1 Data of the top dynamic range of OTDR

* code length 2'M ** cooled by liquid N2

Approach Technology

Wavelength

A

Pulse width T

(us)

Band- width

B

(MHz)

Input optical

power Pio

(dBa)

Mioiaua

detectable

powere

Pd

(dBa)

Postdetection

S/N iaproveaent SN1R

(dB)

Dynaaic

range

L

(dB)

Reference

Year

A

Increase

of input

optical

power

High

Power

Laser

YAG laser 1.32 0.5 0. 3 +30 -75 60 41 (5) 1984

Er***laser 1.55 1 1 +36 -70 " 36 34 (6) 1984

Raian laser 1.59 0.8* 0.3 +34 -70 " 54 35 (7) 1986

Raaan OTDR 1. 06 0.2 2 _ -68 " 39 30 (8) 1984

Correlation Method 1.3 0.1 15 (9) 1984

B

Decrease

of iniaua

detectable

optical

power

Optical preamplifier 1.29 1 1 + 3 -62 30 13.5 (10) 1984

Photon counting

1.32 1 -15 -93 " 34 20 (11) 1981

1.52 1 -26 -75 63 12.5 (12) 1985

Heterodyne detection

1.15 0.8 1.2 - 9 -81 53 32 (13) 1983

1.3 2 -15 57 24 (14) 1984

1.52 1 3 -23 -90 45 10 (15) 1982

fable 2 Advances in OTDR technique other than dynamic range

Subject Technique Feature Reference

Year

Resolution AL

Narrow pulse AL=5aa. pulse width=46ps (16) 1987

Nonlinear detection ALM5*jb. pulse width=0. 12 ps. 300W atA=600 na (17) 1981

Low coherence interferoaetry

AL=38/ia, coherence length=50/ia (18) 1987

AL= ljua, spectral width=130na (19) 1987

Polarization

noise reduction

Acoustooptic deflector t) II /»? -L<1. 18 ( v diffraction efficiency) (3) 1984

Polarization scraabler Tiae averaging by twisting the test fiber (20) 1981

Birefrlngent launching fiber Fluctuation of 0.73 dB is reduced to 0.06 dB (21) 1985

Fading reduction Nonlinear signal averaging Modulus and aul tichannel averaging (22) 1984

User

friendl Iness

Two wavelength aeasureaent A =1.3. 1.5jum

Autoaated wavefora processing Without cursor handling

Receiver -

Dichroic

1 .2~ Mirror

1.7/am

Averager

time

Fig. 1 Raman OTDR technique (8) Fig. 2 Laser-diode amplifier OTDR (10)

5
TIME

SLD P PANDA COUPLER

PRE-AMP

FIBER or WAVEGUIDE

Vn 0. V

«-J

yN XN-1 0Nri X

no. N

ANALYZER

DETECTOR

no. 1 no. 0

Fig. 6 Birefringent launching fiber (21)

Fig. 4 Nonlinear detection (17)

COo

LOCK-IN DVM

AMP

PLOTTER

Fig. 5 Low coherence interferometry (18) ffl 10.

BEAM BIREFRINGENT SINGLE-MODE O

SPLITTER LAUNCHING FIBER FIBER

LASER POLARIZER I i

PULSE WIDTH: 50ns

AVERAGING : 2

17

kii

0 2 4 6 810 12 1416

DISTANCE (km)

2 4 6 8 10

DISTANCE (km)

Fig. 7 Automated waveform processing

6
INVESTIGATION OF OPTICAL COMPONENTS IN MICROMETER RANGE USING AN OTDR SYSTEM

WITH THE BALANCED HETERODYNE DETECTION

R.P. Novak, H.H. Gilgen and R.P. Salath§

Swiss PTT, Technical Centre, 3000 Bern

W. Hodel, P. Beaud, J. Schutz and H.P. Weber

University of Berne, 3000 Berne

Switzerland

1 . Introduction

Optical time domain reflectometry is a well known method for the diagnostics of fiber optics systems. The new small structures such as integrated

optical devices require nevertheless the resolution in the micrometer range.

In our previous work [1] we have reported on the development of an infrared

subpicosecond OTDR system and its first applicability as a diagnostic tool

for the investigations of integrated optical devices. The light source was

delivering pulses of 300 fsec duration and 70 mW average power at the wavelength of 1.3 (Jm. Together with the intensity correlation detection technique the achieved dynamic range was 75 dB with the spatial resolution of

60 pm in air.

In this paper we describe a significant improvement of this system on the

detection side. The detection of the weak backscattered light is performed

with a Mach-Zehnder type interferometer combined with the balanced heterodyne detection, resulting in an improvement of 30 dB compared to the intensity correlation technique.

2. OTDR set-up

Fig. 1 shows the experimental arrangement. The light source consists of a

CW mode-locked Nd:YAG laser, fiber/grating-pair pulse compressor and an infrared dye laser (dye Nr. 26+HFB~) in linear configuration [1] . It delivers

pulses of 300 fsec duration at the repetition rate of 82.00 MHz. The wavelength of the dye laser is tunable between 1.25 and 1.35 |im. The output beam is split by a dielectric beam splitter (BS1) into the Local Oscillator

(LO) part and the Signal Beam (SB). The signal beam is fed through the beam

splitter (BS2) and the lens (L) to the structure under investigation. The

back reflected signal (RS) from the sample passes back to the BS2 and is

directed to the heterodyne balanced detection system consisting of the

third beam splitter (BS3) and the InGaAs JDetectors D-)-D2« The LO beam is

first fed to an acousto optic modulator (AOM) in order to impose a frequency shift needed for the heterodyne detection. In the experiment the AOM

(Bragg cell) is operating at 40.99 MHz. The LO beam passes the AOM twice

being reflected by a 100% mirror (M) mounted on a precision translation

stage (TS) . The total frequency shift is therefore 81.98 MHz. This leads to

a convenient detection of a beat signal at a frequency of 20 kHz. The movement of the mirror M serves to impose a delay for the LO pulses and achieve

in this way the time coincidence with the RS signal (layer sampling)

7
The beam splitter BS3 (50%) is combining at a time two resulting beams - RS

and LO and directs them to two detectors. The electrical beat signals from

these detectors are exactly 180 degrees out of phase due to the phase shift

introduced by the reflection on the beam splitter BS3. Finally the detector

outputs are processed by a differential amplifier (balanced detection) and

spectral analyser.

The induced photocurrents I(T) of the two detectors (D-j, D2

) are given by

the sum of the two Gaussian bandwidth limited pulses:

2 2 2

% T Av

PSB + pLO , „ / ~ r i I±(T) <* ± 2 / PSB PLQ • e 4 In 2 cos [2 % ( vt + vQ

T) + §J 1)

where Plq and PSB are the corresponding powers of the LO and SB at the beat

frequency, Av the spectral width, vQ the center frequency, v the beat frequency of 20 kHz, T the delay and $ the phase between two beams [2]

If the splitting ratio of the beam splitter BS3 is exactly 1/2, the difference of the two photocurrents I+ and I_ will be given by the doubled interference term only. A small difference of this ratio can be additionally

adjusted by two polarisers P1 , P2.

The heterodyne detection of very short pulses allows a very high spatial

resolution.

Experimental results

a) Linearity and sensitivity:

Fig. 2 illustrates the sensitivity and linearity of the system. The

sample is a BK7 mirror substrate of 3 mm thickness. Two highest peaks

(-14 dB) correspond to 4% reflections of the front (F) and back (B) surface. Further, three echos (E1-E3) can be identified resulting in a sensitivity of more than 100 dB. All peaks lie on a straight line demonstrating the linearity of the system over the whole range.

b) Two reflections, small separation:

Fig. 3 illustrates the case, where the separation of two reflections was

smaller than the pulse width in the material. The investigated pellicle

had the thickness of 6.4 |Jm. This leads to the interferences between the

pulses reflected from the front and the back surfaces. These interference effects combined with the tunability of the dye-laser allows to determine such small separations. It can be shown that the optical separation dopt is given by:

Amax • Amin

dopt =

: : ; 4 ( Amax - Amin

)

where Amax and Amin are the wavelength where the maximum positive (one

peak) or maximum negative (two peaks) interference occurs, respectively.

The tuning range (100 nm) allows therefore to resolve two reflections

with a separation of less than 10 (Jm.

8
c) Measurement on Integrated Optical Devices:

Fig. 4 shows an example of a measurement on an integrated optical device. The structure is a rib-waveguide Mach-Zehnder interferometer on GaAs

substrate. Different peaks corresponding to the entrance face (F) -14 dB,

end face (B) - 41 dB, waveguide splitting (Y) -58 dB and echo of the

back-end face (E) -88 dB/ can be detected. Assuming the losses to be distributed evenly, an average loss coefficient a can be calculated from

1 P

a = - In (% • RB

• —

)

2L - Pb

where L is the length of the waveguide and Rp and RB are the reflection

coefficients of the front and end face, respectively. The power ratio

Pjj/Pg of the first echo (E) and back reflex (B) can be taken from the

figure.

4. Conclusions

A new OTDR system using an infrared (1250 - 1350 nm) ultrashort (300 fsec)

pulse source in conjunction with balanced heterodyne detection has been

developed. The resolution is about 60 Mm in air with the possibility to resolve two scattering centres separated by less than 10 pm. The dynamic range is better than 100 dB. The applicability of the system for the investigation of integrated optical devices has been shown. The work is in progress and more examples (optical connectors, splices, couplers etc) will be

presented at the symposium.

5) References

[1] R.P. Novak, H.H. Gilgen, R.P. Salathe, B. Zysset, P. Beaud, and H.P.

Weber

Time domaine reflectometer in the micrometer range. Proceedings of ECI0

87, Glasgow.

[2] R. Stierlin, R. Battig, P.-D. Henchoz and H.P. Weber.

Excess-noise suppression in a fibre optics balanced detection system.

Optical and Quantum Electronics 18, pp. 445-454, 1986.

9
light source 1250 - 1350nm

300 (sec P AV - TOmW

AOM

BS1 BS2

BS3 •

•P1

P2 i

D2

Fig.1 OTDR set-up in micrometer range

SB

RS

01

SAMPLE

OKI. Ampl.

Spectral

Analyser

0.03

0 02

0 01

OS 1 15

mirror translation (cm)

F]g.2 Linearity and sensitivity

1

1

1

1

'

W\ Ja)Mo«.negotive

N interference

— '/

1 V /

1

*/

1 1 f

' \

if 'If'

'

I

1 I J

' v

1 \ /

' if 1 1 /

• '/ 1 \ /

' '/ ' i I

• '/ 1 I /

' if » 1 i i '/ 1 1 /

*

il * \ / i s > 1

1 , >J,

\ : 1 26 6pm

Uin

*

V * \ \ \ » \ X \ \ \i v

1 A

•20

- -(.0 m

3

|< -60 M

QCe

° -80

100

T" 3 JL a

a TT a

J I

I MUn„

1 2

mirror tronflotion lem)

I 2

ao

ti 5

-50 0 50

mirror translation |jjm|

3b)Uox. positive interference

\„ = 1 329>jm Uax

Fig. 4 Reflections from a Mach-Zehnder Interferometer

•50 o 50

mirror translation |pm|

Rg.3 Two reflections. Smoll separation.

10
155 km LONG-RANGE OTDR MEASUREMENT AT 1.55 pm WAVELENGTH REGION

Y. NAMIHIRA, Y. HORIUCHI, and H. WAKABAYASHI,

KDD Meguro Research & Development Laboratories/

2-1-23/ NakamegurO/ Meguro-ku, Tokyo, 153, Japan.

T. OSHIMI/ K. KITAGAWA/ and T. OOKA

Anritsu Corporation, Atsugi, Kanagawa, 243 Japan.

INTRODUCTION : Since the silica-based single-mode fibers

(SMFs) have their minimum loss in the 1.55 um wavelegnth region

[1], 1.55 um range, long-haul optical fiber submarine cable

systems have been developed [2]. In optical fiber systems, the

repeater spacing is expected to be over a 120 km ; therefore,

practical long-range, optical time-domain ref lectometry (OTDR)

is indispensable for the measurement of the loss distributions

and the fault location along such long optical fiber cables.

Untill now, to improve the measuring system dynamic range,

several techniques have been reported, such as using a high

power Nd : YAG laser [3] and cooling a Ge-PIN photodiode to 77 k

(-196 °C) with liquid nitrogen [4] . Other techniques using a

semiconductor laser (LD) and a Ge-avalanche photodiode (APD)

have been widely used, however ; it is difficult to achieve any

improvement in the dynamic range because of the large dark

current of a Ge-APD. Accordingly, a InGaAs/InP-APD [5] with a

small dark current promises to achieve a large dynamic range

OTDR in 1.55 um wavelength region. So far, in 1.55 pm wavelength

range OTDR using a cooled InGaAs/InP-APD with a peltier's

thermodevice, a Rayleigh backscattered signal over a 64 km SMF

with fiber loss of 0.21 dB/km and a one-way backscattered

dynamic range of around 24 dB have been reported [6], Recently,

using high power LD and InGaAs/InP-APD, a 147 km OTDR measurement and a one way backscattered dynamic range of 29.3 dB have

been presented [7] . However, there are few reports on OTDR

11
with a one-way backscattered dynamic range of more than 30 dB.

This letter presents the first measurement results of a

Rayleigh backscattered signal and a Fresnel reflection (with

mirror surface) over a 154.80 km and a 225.45 km SMFs with

fiber loss of 0.187 dB/km by 1.55 urn long-range OTDR using a

InGaAs/InP-APD with a small dark current, a high-power LD and a low noise receiver [8]

A one-way backscattered dynamic range of 32.97 dB at room

temperature has been successfully achieved.

RESULTS AND DISCUSSIONS : To improve the backscattered dynamic

range of the OTDR, signal/noise ratio (SNR) and receiver

sensitivity, the following items were examined : (a) A InGaAs/InP-APD with a small dark current (~8 nA) was used.

(b) A high-power LD module (^v+ 14.8 dBm) was used.

(c) To increase the multiplication factor M of InGaAs/InP-APD,

M was set to 40

(d) The noise floor level of the receiver was reduced to improve

SNR.

(e) The one-way backscatterd dynamic range of the OTDR using

InGaAs/InP-APD was increased by + 9.8 dB comparing with

that using Ge-APD.

Table 1 shows the comparison of the Ge-APD with the InGaAs/

InP-APD parameters at room temperature used in the experiment.

From Table 1, it can be seen that the dark current 1^ of the

InGaAs/InP-APD has a small value of 8 nA (at 0.9Vg) [5], and

when the multiplication factor increases to M = 40, the value

is greater than that of Ge-APD (M = 10). A buried heterostructure InGaAsP/InP-LD at 1.548 urn wavelength is driven at an

effective peak power of + 14.8 dBm (~ 30 mW) at the pigtail

fiber output end. The pulse width is variable and the maximum

width is 10 us. The long-range OTDR parameters are shown in

5 Table 2. The averaging time was 3.8 x 10 times, and the

receiver sensitivity improvement in optical power level was

about 12.4 dB.

Fig.l shows the OTDR measuring system. Fig. 2 represents the

Rayleigh backscattered signal and Fresnel refrection (with

mirror surface) along a 154.80 km and a 225.45 km low-loss pure

silica-core SMFs [1] with a total fiber loss at 1.55 urn of

0.187 dB/km as measured by OTDR using InGaAs/InP-APD.

12
Fig. 3 indicates the one-way backscattered dynamic range

against fiber length for the pure silica-core SMF and the

segmented-core, dispersion shifted fiber (DSF) [9] . The fiber

parameters are shown in Table 3. In Table 3 the group index N^

was 1.462 for the pure silica-core SMF and 1.475 for the

segmented-core DSF [7] . These values were obtained from the

relation between the fiber mechanical length and the optical

group delay time measured by the phase shift technigue.

It was confirmed from Fig. 3 that the one-way backscattered

dynamic range of OTDR was about 32.97 dB. Also, it can be seen

that the difference between the fiber loss measurement results by

the OTDR and the cutback method for the pure silica-core SMF and

the segmented-core DSF are 0.008 dB/km and 0.009 dB/km,

respectively. From these results it is found that loss

measurement by OTDR is in good agreement with that of the cutback

method.

CONCLUSIONS : A Rayleigh backscattered signal and Fresnel

reflection (with mirror surface) over a 154.80 km and a 225.45 km

pure silica-core SMF with fiber loss of 0.187 dB/km was

successfully observed by long-range OTDR using a 1.55 urn high-power LD and InGaAs/InP-APD with a small dark current at

room temperature. Long range 1.55 jim wavelength region OTDR with

one-way backscattered dynamic range of 32.97 dB at room

temperature was developed.

ACKNOWLEDGEMENT : The Authors would like to Dr. T. Muratani and

Dr. T. Yamamoto of KDD Meguro R&D Laboratories for their

encouragement

REFERENCES :

[1] Y. Namihira et al., : Electron. Lett., 23, pp. 963-964, 1987.

[2] S. Yamamoto et al.,: Suboptic '86, France, Communication

13.2, pp. 369-376, 1986.

[3] M. Nakazawa : Electron. Lett., 17, pp. 783-785, 1981.

[4] M. Fujise et al.,: Electron. Lett., 6, pp. 232-233, 1984.

[5] Y. Matsushima et al., : IEEE Electron Device Letters, vol.

EDL-2, No. 7, pp. 179-181, 1981.

[6] K. Shimizu et al., : IEICE, National Convention Record,

vol.70, No. 4, pp. 294-296, 1987.

[7] Y. Namihira et al . , :Electron. Lett., 23, pp . 1219-1220 , 1987.

[8] K. Okada et al., : Electron. Lett., 16, pp. 629-630, 1980.

[9] T. D. Croft et al.,: IEEE, vol.LT-3, No. 5, pp. 931-934, 1985.

13
Test Optical Fiber

LD Driver Hi-Power

Pulse ILD AO

Switch

Timing

Generator AO

Driver InGaAs/ InP-APD

0000

Amplifier

A/D

Averager CPU CRT

Fig.l Long-range OTDR measuring system.

Table 1 Comparison of Ge-APD with InGaAs/InP-APD parameters

at room temperature (~300 K)

Ge-APD InGaAs/InP-APD

Dynamic range

[One way]., (<JB) ~23.13 ~32.97

Breakdown voltage VB , (V) ~33.7 ~79.9

Multiplication

factor, M ~10 ~40

Dark current

[at 0.9VB ] , ~200

Id- (nA) ~8

Excess noise

figure, x ~1.0 ~0.7

Quantum

efficiency, (%) ~ 70 ~ 93

Fig. 2 Rayleigh backscattered signal and Fresnel reflection (with

mirror surface) along a 154.80 km and a 225.45 km pure

silica-core SMFs with fiber loss 0.187 dB/km at 1.55 urn. Fiber loss : 0.195 dB/km (OTDR), 0.187 dB/km (cutback) Vertical scale : 4.0 dB/div, horizontal scale : 8.2 km/div

Pure Silica-Core SMF

180.7:4 km 225.45 km

I 1

80 100 120 140 160 180 200 220 240 260 Fiber Length (km)

Fig. 3 One-way backscattered dynamic range against fiber length

in the case of pure silica-core SMF and segmented-core DSF.

(a) Pure silica-core SMF

(b) Segmented-core DSF

:~0.195 dB/km (OTDR),

-0.187 dB/km (cutback),

: -0.228 dB/km (OTDR),

~0.219 dB/km (cutback)

Table 2 Long-range OTDR parameters

Table 3 Fiber parameters at 1.55 urn wavelength

(a) LD Wavelength Pulse width

Pulse rept. Effective

rate peak power

(within fibres)

Pure silica -core SMF Segmented

-core DSF

(pm) W (us) Tp (ms) P0 (mW) Total fibre length (km) ~ 274.24

[Total fibre loss, dB] [51.42] ~ 193.07

[42.33] ~1.548 ~10 ~ 3.28 ~ + 30 (305 Hz ) (+ 14.8 dBm) Fibre loss (including

splice loss, dB/kra) ~0.187 ~ 0.219

(b) Load resistance AMP R (k ) Band width

B '(MH2 ) [Averaged fibre loss, dB/km] [0.185] [0.207]

Refractive index ~ + 0.87 ~ 10 ~ 3.5 difference, i\) ~ - 0.36 ~ + 0.27

<c) Sampling period

Averager S (jis) Averaged number

N Group index

Ng ~ 1.4 62 ~ 1.475 ~ 1-6 ~ 3.8 x 10 5

14
BIDIRECTIONAL OTDR MEASUREMENTS UTILIZING AN IMPROVED

FOLDED-PATH TECHNIQUE

E. L. Buck land and M. Nishimura

Sumitomo Electric Fiber Optics Corporation

Research Triangle Park, NC 27709 USA

INTRODUCTION

For accurate attenuation and defect analysis using Optical Time Domain

Ref 1 ectometry (OTDR), averaging of measurements taken in both directions

through a fiber is necessary. The use of a mirrored surface at the far end

of a fiber to facilitate a simultaneous two-way measurement has been proposed

[1,2]. However, a noise contribution arising from the finite reflectivity at

the fiber-OTDR interface can significantly degrade the results of attenuation

measurements in the mirrored trace. In this paper, we expand on the theory

of the "folded-path" OTDR and present a practical system, including a

correction algorithm, for significantly improving the performance of the twoway OTDR measurement technique.

THEORY

The theory of single-mode fiber backscatter has been thoroughly developed

[3,4]. The time dependent signal detected by the OTDR is a function of the

total attenuation scattered and confined in the counterpropagat 1 ng fundamental

mode. Material and structural variations along the fiber length result in

scattering factor variations which lead to the bidirectional ity phenomena in

OTDR measurements.

With the folded path OTDR, the impulse is to treat the mirrored trace as an

ideal representation of the reverse fiber signature. This is inadequate,

however. Though the reflectivity at the fiber-OTDR interface is very small,

it is not negligible in the analysis of the mirrored signal. As shown in

Figure 1, additional contributions arise due to the finite reflectivity of the

input end of the fiber.

By analyzing various trace components one can effectively eliminate this

unacceptable signal contribution. Figure 2 shows a typical trace of three

concatenated fibers through a time t=6*L/Vg, where L is the total link length

and Vg is the propagation group velocity. The forward and initial mirrored

trace (through t=4*L/Vg) are represented by the following equations:

Eqn. 1. 0 < t i 2 L/Vg (Tl):

P(t) = Eo Vg e T B(z)

15
Eqn. 2. 2 L/Vg <tU L/Vg (T2)

P(t) = Eo Vg e T R

Q [ e T B(z') + 2e T B(z) —^- J

-2cx 7' -2a^z

Eo is the input signal, is the total attenuation contribution, z and z' are

positions along the fiber axis as measured from the input and the far end

mirror, respectively. B(z) is the local scattering factor. R

g

and R. are the

reflectivities of the mirrored end and the fiber-OTDR interface, respectively.

Similar equations are easily derived for greater times t by recognizing the

effect of multiple reflections.

The mirrored fiber trace of interest is the z' component of Eqn. 2. This

signal is obtained from the measured trace by subtracting out an attenuated

replication of the forward trace (Tl), appropriately time correlated. Letting

DO be the total link attenuation to the beginning of T2, Dl be the return loss

between Tl and T2 and 02 be the return loss between T2 and T3, the

attenuation factor for the forward trace contamination of T2 is

Eqn. 4 -10 log

[

2e R

e

Rj

J = DO + (D2 - Dl) + 6

where 6 is a small error term dependent on the scattering factors B(0) and

B(L), and the end reflectivities.

MEASUREMENT AND ANALYSIS

The measurement system consists of the test link spliced between two

reference fibers , Rl and R2 as shown in Figure 1. Rl and R2 are

depressed cladding single-mode fibers, chosen long enough to suppress

effects of Fresnel reflection detector saturation. Additionally, Rl and R2

have the same mode field diameter, effectively eliminating the error term in

Eqn. 4. A thin layer of gold has been sputtered onto the end of R2 to

provide a mirror surface having greater than 93% reflectivity, R

g

. R2 is a

permanent system component, so that R is a constant for all measurements.

Backscatter measurements were performed using a commercially available

OTDR, utilizing a masking function to further minimize effects of detector

saturation. The large peak at the end of Tl in Figure 2 is the mirror

reflection from the end of the termination fiber. Similarly, the reflected peak

at the end of T2 is caused by the non-zero reflectivity, R., at the input end

of the system. Figure 3 shows an enlargement of T2 from the previous graph

before and after application of the correction algorithm

16
To verify the improvement in measurement accuracy using this correction

routine , 12 cabled depressed cladding single-mode fibers were measured

using the described system. The splices to Rl and R2 were mechanical Vgroove splices with index matching liquid. Attenuation measurements from the

averaged bidirectional OTDR traces were compared with standard cut-back

method results at 1.3jjm and 1.55jjm before and after correction. The results

are shown in Table 1. Recognition of the input reflectivity R. allows for

significant improvement of attenuation measurements.

1.30jjm 1.55jjm ff x <r x o"

| uncorrected trace

|

1 corrected trace

! .036 .032

-.008 .015

.031 .036

j

.011 .017 I

Table 1. Attenuation error: OTDR - cutback (dB/km)

An additional advantage of designing the system with known reference fibers

is the ability to measure mode field diameter (MFD) directly with the OTDR.

With the MFD of Rl and R2 measured by the far field pattern (FFP) method

[5], the MFD at either end of the test fibers were calculated directly from

the splice losses SI and S2 (fig. 1). Figure 4 shows the excellent agreement

achievable between the FFP technique and the simultaneous two-way OTDR

measurement at both 1.3jjm and 1.5jjm. The repeatability of the MFD

measurement using this folded-path technique was measured to have a

standard deviation of <r = 0.1 ^im or better at both wavelengths.

CONCLUSIONS

We have presented a practical system for performing simultaneous bidirectional

OTDR measurements using a reflective termination fiber. A correction

algorithm has been introduced to account for the finite input reflectivity for

improved attenuation measurements.. We have also demonstrated the

feasibility of measuring fiber MFD in the same measurement, greatly

increasing the utility of the OTDR for routine product evaluation or field use.

REFERENCES

1] L. Faltin, J. Opt. Commun., Vol 9, No. 1, 1988, pp. 24-6.

2] M.P. Gold and A.H. Hartog, Proc. 10th ECOC, Stuttgart, 1984, pp. 128-9.

3] D. Philen et.al., IEEE Trans., Vol. MTT-30, No. 10, 1982, pp. 1487-96.

4] M.E. Fermann et.al., J. of Lightwave Technology, Vol. LT-6, No. 4,

1988, pp. 545-51.

5] E.L. Buckland and M. Nishimura, Electron. Lett., Vol. 21, No. 24, 1985,

PP. 1149-51

17
FIGURE 1. Measurement system and optical signal

paths. Tl is the backscattered signal from a point z in the forward trace. T2 shows

possible paths for a scattered signal

arriving at the detector after an additional

round-trip through the fiber. The desired

reverse trace is from the signal at z'.

FIGURE 2. Folded-path OTDR trace through three

round-trip traverses of the input signal.

FIGURE 3. Reverse OTDR signature a) before

correction for the reflectivity Rl and

b) after correction.

FIGURE 4. Mode field diameter measurements:

Far field pattern vs. bidirectional OTDR

at 1300nm and 1550nm.

Mode Field Diameter (pm): FFP

Fig.4

18
OTDR EXPERIMENT FOR GLASS OPTICAL WAVEGUIDES BY USING

INTERFEROMETRIC TECHNIQUE WITH 1 . 3 jam SLD

K. Takada, H. Nagai*, N. Takato, and J. Noda

NTT Opto-Electronics Laboratories,

Nippon Telegraph and Telephone Corporation

Tokai , Ibaraki-ken, 319-11, Japan

Tel. 0292-87-7515, Fax. 0292-87-7877

* Morino-sato, Atsugi-shi , 3-1 , Kanagawa-ken , 243-01, Japan

Tel. 0462-40-2751, Fax. 0462-40-7863

Abstract

First demonstration of OTDR experiments on the glass optical

waveguides is presented by using a newly proposed OTDR system

based on a Michelson interferometer. To obtain high spatial

resolution less than 0.1 mm, an 1.3 yam-wavelength

superluminescent diode is adopted. Scattering centers produced by

waveguide irregularity are observed clearly in silica based

waveguides

Introduction

Optical time domain reflectometer s (OTDRs) are now widely

used for testing optical fiber cables. The spatial resolution of

the conventional OTDRs using optical short pulses is relatively

low, about lm, because of limit in pulse width and the frequency

characteristics of the electrical circuits. Therefore, the

conventional OTDRs are not applicable for characterizing

miniaturized optical waveguides. In this paper, we present a new

type of the OTDR based on the inter f erometr i c technique and

discuss experimental results on application to silica-based

optical waveguides.

Experimental setup

The proposed OTDR system is shown schematically in Fig. l. 1 ^ The system is composed of the superluminescent diode (SLD), a

Mach-Zehnder interferometer and a Michelson interferometer. To

achieve high spatial resolution, the incoherent and high power

SLD is used as a light source. The SLD light oscillating at the

center wavelength of 1.3 yum is launched into the fiber-optic

Mach-Zehnder interferometer, i.e., into one arm C^ of a

polarization-maintaining PANDA fiber coupler. 2 ^ The launched

19
light is divided into two arms C 3 and C^. Light in the arm Cg

reflects at a mirror MQ

, and the reflected light is used as the

reference light in the interferometer. The fiber arm C

4 is

connected to a test waveguide through a cylindrical

piezoelectric phase modulator (PZT MOD) which modulates the

propagating light beam at a frequency of 10 kHz. The Mach-Zehnder

interferometer with the fiber coupler is necessary for coupling

the light into the test waveguide without excess loss, and for

superimposing the backscattered light with the reference light.

Here the polarization maintaining fiber coupler ensures the

stable performance little dependent on the polarization noise.

The superimposed light at the coupler is guided into the

bulk-type Michelson interferometer through the arm C

2

• The output

signal from the Michelson interferometer is detected by a lock-in

amplifier at the frequency of f+ Af, where f is the modulation

frequency and /\ f is a Doppler shift caused by the scanning of a

mirror M2

Principle

When a light propagates in a waveguide, a fraction of the

light is back-scattered at arbitrary positions in the waveguide.

The optical path difference A£ between the particular two

reflected waves is twice the distance AS between the two

scattering points multiplied by the refractive index n

c

of the

waveguide material, i.e., j\H =2/dSn c

. The two waves become

independent of each other when is longer than the coherence

length J!c of the light source. These two waves interfere

independently with the reference light. To make these light waves

interfere, the Michelson interferometer is necessary for giving

the corresponding delay.

In the Michelson interferometer shown in Fig. 1, optical

path P-^ of the first arm is held constant, while optical path P2

of the second arm is variable. When P 2 is varied, the

interference envelopes appear in the Michelson interferometer. As

the reflected light is phase-modulated with f=10 kHz, only the

interference signal between the reference light and the returning

light from the test waveguide can be detected by the lock-in

amplifier. The interference peak produced at a given mirror

position is proportional to "f , where Y is the amplitude of the

returning light at the corresponding back-scattering point. The

relative power at this point is described by T

7 ^ » or

logarithmically, by C=-201og f ( dB )

20
Results

Figure 2 shows the I-L characteristic of the SLD operating

at ^=1.3 /im wavelength. 3 ^ The output light power is 1.2 mW at

1 = 140 mA and its spectrum is almost continuous with §/l =40 nm

spectral width. The output power was very stable and no spectral

changes were observed while the OTDR experiments were made. The

coherence length was £c = A~/£/L =42 /mi, and then the spatial

resolution of the present OTDR was j^./2n

c

=15 /im.

Test silica glass optical waveguides are single-mode

embedded straight waveguides fabricated by a combination of flame

hydrolysis deposition and reactive ion etching processes. 4 ^ The

cross-sectional view of the glass waveguide is shown in Fig. 3.

The core consisting of Ti02 doped Si02 was 8 jum in both width

and height. The buffer and cladding glass layers made of Si02

were 8 yum and 32 ^um in thickness, respectively. The refractive

indices of the buffer and the cladding layers were almost the

same, and were lower by 0.25 % than that of the core.

Figure 4 shows the measured amplitude f

7

vs . the optical path

difference in the interferometer, where the test waveguides are

1.8-cm long SiC^ glass optical waveguides. These two optical

waveguides were fabricated under different process conditions,

although their waveguide parameters were the same. Both ends of

the test waveguides were immersed in matching oil to reduce the

end reflection. For the single-mode fiber, the reflected power at

both ends were reduced to lower than -70 dB relative to the

propagating light power. 1 ) However, the reflected power did not

decrease to lower than -40 dB for the glass optical waveguides

with the same matching oil. This difference may be due to the

refractive index difference between the waveguides and the

matching oil and/or due to imperfect end preparation of the

waveguides. Thus the relatively-large end reflections were

observed for these glass waveguides.

In the single mode fibers there are no scattering centers

except for Rayleigh scattering centers. Correspondingly, no peaks

were observed at the detection range of J

7 =10~ 5 when a short

piece of a single mode fiber was used as a test waveguide.

However, a few scattering centers were observed for one glass

optical waveguide as shown in Fig. 4(a), while no scattering

centers were observed for another as shown in Fig. 4(b).

Therefore, it is clarified that the scattering centers are

produced in the silica based optical waveguides under slightly

different process conditions. The detection level is set at one

order of magnitude higher than Rayleigh-backscattered signal

21
level. Therefore, thes

called Mie scattering,

inhomogenei ty in the

consolidation

e observed peaks may

The origins of the

waveguides produced

be produced by the soscattering centers are

during etching and/or

Conclusion

We have proposed a new type of OTDR system for

characterizing miniaturized optical waveguides by using

interferometric techniques with a 1.3 ^um SLD. Scattering centers

in the waveguides were made clear. Therefore, the system is

useful for finding the scattering centers and for clarifying loss

mechanism in the optical waveguides.

Acknowledgment

The authors would like to express their sincere thanks to I.

Yokohama for supplying the PANDA coupler and to E. Noguchi fo

fabricating 1.3-yum SLDs. They also thank to N. Uchida for his

continuous encouragement.

References

1) K. Takada et al . , Appl. Opt. vol. 26, pp. 1603-1606 (1987).

2) I. Yokohama et al ., Electron . Lett. vol. 22, pp . 929-930 ( 1986 )

3) H. Nagai et al ., submitted to Electron. Lett.

4) N. Takato et al ., Electron. Lett, vol.22, pp. 321-322 (1986).

SLD PANDA COUPLER

r~~

M,

HM

STAGE

PRE-AMP

InGaAs-PD

IS

FIBER or WAVEGUIDE

LOCK-IN DVM

AMP

PLOTTER

Fig. 1. Experimental setup

22
Fig. 2. I-L characteristic of SLD . Fig. 3. Cross sectional view

of Si0 9 glass waveguide.

POSITION AS (mm) POSITION AS (mm)

OPTICAL PATH DIFFERENCE AS (mm) OPTICAL PATH DIFFERENCE AC (mm)

Fig. 4(a). Amplitude J

7 vs . optical

path difference for 1.8cm-long

SiC-2 glass waveguide.

Fig. 4(b). Amplitude J

7

vs . optical

path difference for 1.8cm-long

S1O2 glass waveguide.

23

Reflectance Measurement in Lightwave Systems: A Comparison of Various

Techniques

S. K. Das, A. F. Judy, G. M. Alameel, R. M. Jopson, T. F. Adda

AT&T Bell Laboratories

1. Introduction

The measurement of the reflectance of fiber optic components is increasing in importance as operating

systems push the performance limits of fiber. For example, the performance of semiconductor lasers,

especially single-line ones, can be either degraded or improved by optical feedback'1 "2"3

) from nearby

connectors or from fiber backscatter.'4

' Both coherent and direct detection systems can be effected by

intensity noise caused by multi-path interference between two or more large reflections

5

' Likewise bidirectional systems'6

! and analog systems'7

' may be sensitive to large reflections.

This paper overviews the principle methods for measuring reflections and discusses their applications and

limitations. It also presents a comparative "round-robin" study between five different measurements at three

AT&T Bell Laboratories locations: Whippany, Crawford Hill and Atlanta. Close agreement has been

obtained between these measurements and estimates are derived for the repeatability and accuracy of the

various test methods. We conclude with identification of problem areas and recommendations.

2. Types of Optical Reflection Measurements

The typical method for measuring reflections uses a separate light source and detector connected by a

coupler. (Fig. 1(a)) The coherence length of the source must be much shorter than the distance between the

components so that coherent interference will not occur. Also the system, especially the coupler, must be free

of polarization sensitivity. (We note that the evaluation of polarization effects is the least developed in the

literature.) Similarly, mode dependent coupling must be negligible in multimode couplers that are used to measure reflectance in multimode fiber systems.

The type of modulation of the light source can be used to divide optical reflectometry into three subtypes:

optical continuous wave reflectometry (OCWR) which is often called the "coupler method", optical time

domain reflectometry (OTDR), and optical frequency domain reflectometry (OFDR). The first two are the

most commonly used and will be discussed in detail below. OFDR, which is also called optical network

analysis, is not widely used at the present. It may become popular for analog systems applications since it operates by frequency modulation of a rf/microwave subcarrier onto the light source and detecting (and

displaying) the magnitude and phase of the reflectance as a function of the modulation frequency.'8

' However

we will not treat it extensively here.

2.1 OCWR: Description

OCWR is perhaps the simplest and most developed of the reflectometry methods. At least two

measurement standards have been proposed to standardize it.'

9 "10

' However both require an inconvenient

making-and-breaking of fusion splices to perform the test. Here we describe a more general procedure and

theory which is inclusive of the original methods while eliminating this restriction.

The OCWR method functions by generating an essentially unmodulated, CW signal. (A very low

frequency modulation for lock-in detection is equivalent.) Thus the detector measures the sum of the power

from: (1) the coupler crosstalk, (2) the reflections from all the components-under-test, and (3) any spurious

reflections. The following procedure describes the measurements and calculations needed to separate these

contributions.

25
2.2 OCWR: Theory

The power transfer between the four ports of a bidirectional 2x2 coupler, can be represented by a symmetric 4x4 matrix [tj\ . For an application such as measuring unknown reflectance values, port 4 is unused

and hence terminated inside the coupler with a suitable index matching epoxy. This is shown in Fig. 1 (a).

Port 3 is connected to the test system; ports 1 and 2 are connected to a light source (preferably a laser) TX

and a power meter RX, respectively. Fig. 1 (b) shows a test connector Cx connected to port 3 of the coupler

via an intermediate connection C,-. Assuming that no reflection is coherent with any other, the received power

PR i at RX is given by

Pr\ = Pt { nts + iisv&Ri + tiun&niR, + iiMnihlR-m + vi^R* ] (i)

where P^laser power output, P,=reflection from connection C,-, i?I =reflection from connection Cx , P4 =residual reflection of fiber termination at port 4, Py^residual reflection from termination Tl, f?,-=power

coupling efficiency of C,-, and ?? I =power coupling efficiency of Cx

. The test connector Cx

is opened and the

leading connector/splice is terminated as shown in Fig. 1 (c). (Alternately, the fiber before test connection Cx may be looped in a small diameter to attenuate the reflection from Cx .) The received power at the RX is

Pr2 = Pt [ *?i2 + VizVwPi + VnnwtfPTz + VuVwRi] (2)

where Pj-2

is the residual reflection from end termination T2. The power emerging from T2 is given by

P3 = PTVizVi • Subtracting (2) from (l) and rearranging the terms one gets

R, = R',+ [RT(l-t,i)} , (3)

where the raw reflectance is given by

R

, _ Pr\~Pr2

PzViVw

and Rt=Rti=Rt2 has been assumed. (This assumption has been verified, to a limited extent, in experiments

involving a given batch of connectors polished in a similar fashion.) From (3) it is clear that the crosstalk nl2 and the reflections P4 or P,- have no influence in the measurement of Rx . Hence, index matching C,- is

unnecessary, unless P,- is so high that it masks the sensitivity of the measurement of PR1 and PR2 - What is important, however, is the end reflection Pj- which has traditionally been ignored. R'x is an easily measurable

quantity and so is nx

. The "correction" ARX in dB is given by

ARX {dB) = 10 log Rx - 10 log R'x = 10 log [ 1 +

R^1 ~r'^

] ( 4) Rx

and is plotted in Fig. 2. It is found that for a test connection loss £,=-10 log 7/,=0.5 dB, 10 log Pr=-30 dB

and 10 log P'^-40 dB, the correction ARX (dB) is « 5 dB. The correct reflectance is then -40+5=-35 dB.

2.2.1 Example : OCWR Measurement of Fresnel References. It will be shown that the OTDR methods for

measuring reflectances rely on either of the two references, Rayleigh backscatter or a glass-to-air Fresnel

reflection. The OCWR provides a simple and accurate means (±0.5 dB) to measure Fresnel reflections. This

is done as follows. In Fig. 3 (a), jumper BC has a reference Fresnel reflection at the end C. With B connected

to port 3 of the coupler one obtains

Pr\ = Pt [ Vl2 + fllZ^Ri + VlsVsztf&F + '/l4 ??42^4 ] ( 5)

where RF is the reference Fresnel reflection. Now the end C is index matched as in Fig. 3 (b), and the power

26
PR2 at RX is:

Pr2 = Pt [ V12 + ViiVszPi + VisV32^RT + Vurji2 RA ] (6)

where Rx is the residual reflection from the termination. Subtracting (6) from (5), and using (3) we get

Rp = R'x + Rf. Usually, the Fresnel reflection value is in the order of -14 dB, whereas Rx (dB) is more like

-35 dB. Hence, little error is incurred if one assumes Rf — Rx

.

2.3 OTDR: Description

The OTDR transmits narrow pulses into a fiber and displays the reflected signal versus time or,

equivalently, distance. If the pulse width is narrow enough, the reflections from the components-under-test

are separated from each other as well as from the coupler. Thus the calculations involved for the OCWR

method are eliminated.

The use and specification of OTDR testing has been well documented for almost every application!11 '

except for reflection testing so we will briefly summarize its use here. Since all OTDRs display the reflected

signal on a relative scale, one must first calibrate with a known reference reflection. The most readily

available reference is the Rayleigh backscatter from the fiber itself which is:

Rbt =^PWat S

where v

g is the group velocity; PW is the temporal pulse width of the equivalent rectangular pulse; a, is the

fiber attenuation due to Rayleigh scattering; S is the proportion of backscatter that is captured by the fiber

(which is a function of the fiber's spot size). The uncertainty in calculating Rb, is typically a few tenths of a dB due to fiber parameter uncertainties.

For wide pulse widths, Rbe may approach the reflectance value of the component-under-test and the OTDR

will display the power sum of the two. R^, must be subtracted to obtain the correct reading.! 12

! If very narrow

pulse widths are used, R^ becomes too small to measure. In this case one can also calibrate with a glass-to-air

Fresnel reference (which can also be used to check the Rf,, value). Caution is advised in that many commercial

OTDRs were designed to measure only small reflections, and may be saturated or nonlinear for large ones.

3. Reflectance Comparison Experiment: (Round-robin)

Twelve test connections were made from pairs of 35-foot single-mode jumper cables which had ST

connectors on either ends. The test connections, numbered from 1 through 12, were not disturbed until all the

reflectance measurements were complete. These were then measured on two different OCWRs and three

different OTDRs. Since the OCWR method usually requires making and breaking the test connection, the five

measurement techniques were performed in the following sequence:

3.1 Whippany: OCWR

The procedure of Sec. 2.2 was followed to measure the reflectance of the 12 test connections. Correction

factors were ignored since Rx was small (<-35 dB).

3.2 Whippany: Millimeter OTDR

The OTDR used here consisted of an assembly of components manufactured by Optoelectronics, Inc. The

laser is modulated by a 0.5 ns pulse, yielding a resolution distance of 5 cm in the fiber. The repeatability of

this system is « 0.3 dB. Since it could not measure Rayleigh backscatter, a rotary splice ferrule polished with

chromium oxide was used as the reference Fresnel reflection; its reflection was determined as -14.7 dB using

27
the procedure described in Sec. 2.2.1.

3.3 Crawford Hill: 50-cm OTDR

This reflectometer has a received pulse length of 5 ns resulting in a resolution of 50 cm. It was calibrated

with a Fresnel reflection. Since the jumpers being tested were longer than 50 cm, it was not necessary to

correct the data for other reflections. However, a correction was needed for the loss of C,- (Fig. 1 (b)). This

loss was determined by comparing the power out of coupler port 3 before connection of the jumper assembly

with the power out of Tl after the connection and subtracting the loss in the test connection Cx . (The loss

value of Cx was taken from Sec. 3.1.)

3.4 Atlanta: 200-m OTDR

This technique uses a 1.3-(Um Anritsu MW98A OTDR with a 2-^s pulse. Normally the test connection, Cx ,

is positioned at the end of a long length of fiber and the OTDR simultaneously displays the fiber backscatter

and the reflection from Cx

. The difference between the two is measured, then by knowledge of the amount of

fiber backscatter, Cx s reflectance is found.

Since the 2-fis pulse has a spatial width of over 400 meters, this method requires that any two reflections

be at least 200 meters apart if they are to be independently visible on the OTDR. An even greater separation

is required to view the fiber backscatter. So several hundred meters of fiber must precede and follow Cx for

accurate measurements.

In the experiment reported herein Cx was the junction of two ST connectors, each on the end of a 35-foot

jumper cable. This resulted in three reflections, [C\, Cx , and C2 as in Fig. 4) which could not be

distinguished from each other. Consequently the following 7 step procedure was used to measure the

reflectance of Cx instead of the usual method:

1. The power, P2 , from C2is measured with a power meter. This is for the later determination of the

insertion loss of the first two ST connections. Fig. 4(a)

2. An approximately 1/8" diameter loop is tied in the fiber that follows Cx. This attenuates the reflection

from C2 to well below Cx s value. The combined reflection from Cx and connector C\ is measured on

the OTDR. This is designated as Rm = R^+nlRx- Fig. 4(b)

3. An approximately 1/8" diameter loop is tied in the fiber that precedes Cx

. This attenuates the reflection

from Cx and C2

. The reflectance R x is measured on the OTDR. Fig. 4(c)

4. Connection Cx is disassembled and the power, Pit from the 1-km test lead is measured. Fig. 4(d)

5. The insertion loss of the two STs is calculated as:

ViVx = (7)

6. Normally Cx would be disassembled and the power measured to determine r\ x as in the preceding two

steps. However disassembling the connectors would have precluded further measurements so instead the

value from Sec. 3.1 was used.

7. Rxis calculated to be:

Rx = ^Jl^. (8 )

where all the parameters on the right hand side are known, ??i being determined from steps 5 and 6.

28
3.5 Atlanta: OCWR

The measurements followed the procedure outlined in Sec. 2.2. Correction factors were not accounted for

since the end reflectance was either comparable to, or lower than, the measured values of ST reflectance.

However, instead of measuring 77,- every time a new jumper connection is made to port 3 of the coupler, the

Atlanta method uses a standard value of L,=-10 log 7?,=0.5 dB, to facilitate rapid reflectance measurements.

This 0.5 dB value for L,- has been substantiated from loss experiment with a large sample size of STs.

4. Results and Discussion

In Table 1 are listed technique-to-technique reflectance variations

Vi = RX{ - JTX (dB)

where t=l to 5 and TTX denotes the mean reflectance of a ST connection over the 5 measurement techniques.

Based on the small values of standard deviation a (p < 1 dB for 9 out of 12 STs), it is clear that both the

OCWR and the OTDR methods yield accurate results. The sum of the five v,s across some rows do not add to

zero due to rounding errors.

When compared to the OTDR data, the OCWR results were slightly more optimistic due to the neglect of

correction factors which account for unwanted reflection from index-matched terminations. In such cases, fiber

loops are shown to offer a simple and effective alternative to index-matching, since the (unwanted) returned

power can be suppressed to the desired level by increasing the number of loops and/or by decreasing the fiber

bend radius within safe limits.

To summarize, the various methods compare well but each has advantages in certain applications. For

instance if one is measuring closely spaced components and wants to avoid the inconvenience of fiberlooping/index- matching, a properly calibrated, high resolution OTDR is superior to some of the other

techniques described in this paper. If on the other hand one wants to measure the total power reflected from

numerous components (including fiber backscatter), the OCWR is the best. Or if one wants to measure the

reflections from widely spaced splices on a long installed fiber route, a wide pulse-width OTDR would be

optimum. Finally, while not evaluated here, OFDRs may be advantageous for measuring the effects of

multiple reflections on the analog transmission systems.

References

1. O. Hirota and Y. Suematsu, IEEE J. of Quantum Electronics, QE-15, p. 142, March 1979

2. R. W. Tkach and A. R. Chraplyvy, J. Lightwave Tech., Vol. LT-4, p. 1655, Nov. 1986

3. W. I. Way and M. M. Choy, J. Lightwave Tech., Vol. 6, No.l p. 100, Jan. 1988

4. A. R. Chraplyvy, D. Marcuse and R. W. Tkach, J. Lightwave Tech., Vol. LT- 4, No.5, p. 555, May 1986

5. J. L. Gimlett et.al., OFC 88, paper WG4

6. P. P. Bohn and S. K. Das, J. Lightwave Tech. Vol. LT-5, No.2, p.243, Feb 1987

7. K. Fujito, et.al., OFC 88, paper THO

8. M. Nazarathy, D. W. Dolfi and S. A. Newton, OFS 88, paper WDD7-1

9. Electronic Industries Association, "Return Loss", FOTP-107 (proposed)

10. Bell Communications Research, TR-TSY-00326, Issue 1, August 1987

11. B. L. Danielson, Applied Optics, Vol. 24, N0.15, p. 2313, Aug. 1,1985

12. A.F. Judy and H.E.S. Neysmith, in Proc. SPIE, 574, Fiber Optic Couplers, Connectors and Splice

Technology II, p. 146 (1985)

29
Fig. 3. (a) Set-up to measure reflectance of the

(c) Fresnel reference Jumper BC has the

Fresnel reference at end C. (b) End C is index-matched.

Fig. 1. (a) Coupler connections in a typical reflectance measurement set-up. (b) Test connector/splice C, connected to port 3 of the coupler via intermediate connection Q. (c) C, is opened and the leading connector/splice terminated.

OTDR

1 ton w 1m H -US-*

Ci C, c 2

u>

(t»

1 1-

-30 dB

1 dB

I R T «-30dB ^ 4 tL«« 0 5 OB

-40 -35 -30

REFLECTANCE R', (dB)

Fig. 2. Correction AR, in dB, as a function of the raw reflectance R, (dB), to account

for the non-zero end reflection from ST

ferrule termination Tl.

OTDR C M c, c }

LOOP

Fig. 4. Set-up to measure reflectance of Cz with

a 200- m OTDR.

Table 1 Technique-to-technique Reflectance Variation (dB)

ConnectioD #

Whippany

OCWR

Whippany

OTDR

Crawford Hill OTDRAtlanta OTDR

Atlanta OCWRa

1 0.3 -0.4 -0.4 -0.2 0.5 0 4

2 1.9 0.3 -0.8 1.8 -3.0 2.0

3 -0.2 -0.5 -0.2 0.5 0.3 0.4

4 0.6 0 4 0 -0.2 -0 9 0 6

5 0 0.4 0.2 0.1 -0.5 0.3

6 -3.3 0.4 1.2 1.6 0.3 1.9

7 -0.1 -1.2 0.5 0 5 0.3 07

8 -0.4 11 1.0 0.3 -2.2 1.4

9 -0.4 -1.1 0.7 1.0 -0.2 0 9

10 -0.4 -0.2 0.3 0.1 0 0 3

11 -0.2 0 -0.1 -0.4 0.9 0 5 12 0.1 0.9 -0.3 0.1 -0.8 0 6

30
Calibration and Use of an Optical Time Domain Reflectometer For Measurements

of Reflectance (Return Loss) of Optical Fiber Components

Paul R. Reitz

AMP, Incorporated

Harrisburg, PA.

Introduction

The typical measurement of return loss (optical source, coupler, and power meter) has

several limitations which can be addressed with the OTDR. Use of the OTDR is not new 1 , but

has not become widespread, partly due to questions about accuracy. In a manner implied by

Judy and Neysmith 2, we report a straightforward calibration method for the OTDR based on the

use of known index oils with a cleaved-end fiber.

Description of the Method

Most workers approached measurement of reflectance with the OTDR by calculating the

backscatter level to enable determination of the relative power produced by a reflectance. Our

method is to calibrate the instrument together with an optical fiber by measuring the reflection

spike amplitude, Ar

, from a length of optical fiber as a function of the index of a medium into

which the output end of the fiber is placed. This data is then used to determine the effective fiber

index, n \

, from which a scale factor K for the OTDR can be determined. Return loss can then

be readily calculated. Once the effective fiber index at the measurement wavelength is known, a

much simpler procedure is sufficient to calibrate other instruments.

Theory

Referring to Figure 1 , the displayed signal amplitude, Pj , on an OTDR trace at point 1 just

prior to the output end of the fiber can be generalized as P! =C! -c2

' Pinc ,k (D

where cj is an instrumentation factor (including return-trip fiber loss), Pjnc is the incident power

propagating at point 1 , and k is the backscatter fraction for the fiber that includes both scattering

loss and backscatter capture. C2

includes the effects of pulse duration, which affects backscatter

without affecting the peak magnitude of the reflected signal. The displayed signal amplitude at

the peak of the reflection spike, P2> is modeled as P2 = Pi + C! • c3

• Pinc • R

(2)

where C3 is a detection system pulse response factor. R is the reflectance of the fiber /medium

interface, given by the reflection equation for normal incidence:

R={[ ni -n2 ]/[ n!+n2 ]}2 (3)

where nj is an effective fiber material index at the measurement wavelength, and n2

is the index

of the medium into which the fiber is terminated.

31
The reflection spike amplitude, Ar

, read from the OTDR display from the base to the peak

of the reflection spike in decibels (10 log vs. 5 log), is given by

Ar =101og[P2 /Pi]

= 10 log [ 1 + K.R] (4)

where K = C3 / c2

• k . The return loss, Lr , (ignoring connection loss), is calculated as Lr = 10 {

- log K + log [ 10 (A

f

/ 10) . 1]} (5)

Experiment

K was determined by measuring Ar for specific values of R calculated by determining

appropriate values for nj and n2 , using a typical single-mode fiber with a cleaved output end.

Previous work has shown that polished ends should not be used. Since an exact value of nj was

not known, Ar was measured for a series of 20 oils covering the range of n from ~ 1.39 to «

1.51, with the fiber sequentially terminated in each oil. Oil indices were translated to the OTDR

wavelength using Cauchy's equation.

A value of nj was iterated until the computed K values were nearly equal for both high

and low n's. The resultant value of K was used to calculate theoretical Ar values using equation

4, which are compared with the measured ones in Figure 2. Figure 3 shows the error as a

function of the computed return loss. The agreement is excellent for all but the largest and

smallest values of Ar. The error at large Ar

is amplitude nonlinearity while the error at small Ar

is both amplitude uncertainty (related to a periodic undulation in the backscatter level, a

characteristic of this OTDR), residual n\ uncertainty, and limitations of the simple model..

The relationship between measured Ar and return loss is shown in Figure 4 for three

different pulse duration settings of the same OTDR; this instrument could measure return loss on

single-mode fiber from -25 dB to at least -60 dB. This range is ideal for the specification

requirements likely to be placed on passive components, as values above -30 dB would typically

constitute failure for low-reflection components. Table 1 summarizes the results for various

fibers and instrument settings. The single-mode data suggest that only small errors would be

introduced with a fiber other than the calibration fiber.

Primary sources ofuncertainty with this method include (1) linearity and noise of the

OTDR amplitude scale, (2) errors in the value and stability of K, (3) end angle on the calibration

fiber, and (4) connection loss to the component to be measured.

The OTDR is readily characterized in terms of the maximum usable Ar value, based on the

departure of the theoretical curves from the measured data, and Ar uncertainty. Care must be

taken to operate the instrument in this region. Typical reproducibility of Ar

is on the order of

0.2 dB (3 sigma) for 5 dB < Ar < 22 dB.

While only two Ar and index data are sufficient, in principle, such a procedure would not

be robust against either calculation or contamination error in an index value. By performing the

calibration with many index oils, the determination of K is quite independent of a single oil index

32
error. The fact that measured and calculated Ar

's fit with so little apparent fluctuation in n

suggests that the extrapolation error is more systematic than random. A systematic error of

0. 5. in n causes < 0. 1 dB error in Ar

.

Figure 5 shows that the reduction of reflectance caused by fiber end angle follows

accepted theory sufficiently well to be correctable for small end angles. The error reduction that

can be achieved depends on error in the measurement of end angle; in many cases the end is

flatter near the center of the fiber.

The loss of the connection with which the component to be measured is attached to the

calibration fiber can be measured with the OTDR, given sufficient fiber length. Higher optical

resolution OTDRs would be helpful, but an optical power meter can also be used.

Comparison with the Coupler Method

Results obtained with the OTDR have been compared to the coupler method on sixteen

multimode components exhibiting return losses ranging from -26 to -32 dB. A mean difference

of 0.61 dB was observed, with the OTDR indicating higher return loss. Return losses have been

measured on mechanical splices in excess of -60 dB.

Conclusions

The OTDR is a viable method for the quantitative measurement of reflectance from optical

components, and will become more attractive as further advances are made in OTDR optical

resolution. Measurement of return loss over the range of 85 -25 to -60 dB with an estimated

accuracy of 55 1 dB is feasible. The calibration method can also be used in the coupler method to

eliminate the need for a known-reflectance "golden" lead.

Acknowledgements

The author would like to thank the following individuals for their contribution: S. J. Harms, oil index calculations and 100/ 140 data; R.H.Rebuck, procedure and single-mode data;

and W. K. Long, comparison data.

References

1 . Methods for Reducing Optical Fiber Connector Reflection Through Use of Index

Matching Materials.

I. Sankawa, T. Satake, and N. Kashima, Review of the Electrical Communications

Laboratories (NTT), Vol. 34, No. 6, 1986.

2. Reflections from Polished Single-mode Fiber Ends. A. F. Judy and H. E. S. Neysmith,

Proceeedings of SPIE, Vol. 574, 1985.

33
Fiber Size / Type K Pulse Width

. t

10/ 125 matched-clad single-mode

780000 342000

160 ns 640 ns

Z.JO US

1.4520

9.5/ 125 matched-clad

single-mode 700000 160 ns 1.4517

62.5/0.275 multimode,

850nm.

6500 160 ns 1.4795

100/ 140 multimode, 850 nm.

20000

6000

20 ns

160 ns 1.4850

TABLE 1 Summary of Results for Various Fibers

Figure 1. Calibration Set-up and OTDR Trace

index, 1300 nm Return Lost, dB

Figure 2. Calculated and Measured Ar Figure 3. Error in Ar vs. Measured Return Loan

FRESNEL REFLECTION POWER VS RETURN LOSS

Figure 4. Ar vs. Pulse Width and Return Loss

34
OTDR MEASUREMENTS OF OPTICAL RETURN LOSS

Felix P. Kapron, Edward A. Thomas, and John W. Peters

Bellcore

445 South St.

Morristown, NJ 07960-1910

Abstract

Precautions are given for using an optical time-domain reflectometer to measure the return

losses of discrete reflections. We present a new method for determining the backscatter level of

the fiber type under test, and show its importance in optical return loss measurements. Some

improvements for OTDR manufacturers are offered for consideration. A new EIA FOTP on

optical return loss via OTDR is warranted for both component and link characterization.

Introduction

Reflections from discrete points along a fiber optic link can degrade transmitter or receiver

performance in high-speed systems [1,2]. To avoid an excessive link power penalty, the

minimum allowed optical return loss (ORL) for elements such as splices and connectors is sometimes specified [3]. In terms of reflectance (captured reflected power Prfl over incident

power Pinc ) of the element, and in positive dBs

ORL = 10 log10 ^P- . (1)

The usual method for measuring ORL utilizes a 2x2 coupler [3,4] and is somewhat tedious. In

contrast, the optical time-domain reflectometer is uniquely qualified to efficiently measure both

the magnitude [5] and location [6] of a sequence of reflections that may require corrective

action in the field. However, since OTDRs are not yet commonly used for this purpose, current

instruments are optimized for attenuation (rather than ORL) measurements.

In this paper we show that the accuracy of the OTDR method is sensitive to the correct

values of pulse duration, pulse height, and fiber backscatter level. We obtain these values by a combination of coupler and OTDR techniques. Finally, we suggest improvements for future

instruments.

Return Loss Using a Coupler

As shown in Figure 1, a source is connected to coupler port o, and a fiber reel is fusionspliced to port c. For a partially reflecting far-end under test we measure powers Pinc out of

the fiber far-end (which is actually slightly smaller than the true incident power); PR

out of

port 6, when the far-end is reflecting; P0 out of port b, when the far-end is non-reflecting.

The reflected power (PR

- P0) has suffered a loss in travelling from the far-end to c to 6.

This loss is measured by first splicing the source to the far-end, and measuring Pgut at port 6.

Then the fiber reel and coupler (but not the splice) are removed by cutback, and Pin is measured. The optical return loss is calculated from

ORL = 10 log,, - %j

f% . (2)

35
detector

(p.)

coupler

L_ if 1 1 fiber

non-reflector

Figure 1 - ORL using a coupler

\ P

i • inc (R=o)

Return Loss Using an OTDR

A reflection from a fiber discontinuity is displayed on an OTDR as a logarithmic replica of

the incident optical pulse, modified by the receiver response and any fiber attenuation and

dispersion. It is superimposed to a height P (in dB) above the fiber backscatter level (Figure

2). Under the assumptions of a "rectangular" optical input pulse of time duration D (in ns)

and negligible pulse broadening, the optical return loss is [7] ORL = B - 10 log 10 [(10

p/5 - 1) D] . (3)

B is the fiber backscatter level in dBs below the power Pinc of a pulse having a 1-ns duration

incident upon the reflective discontinuity. A quantity related to B can be determined [8], but

we measure B by a new method as shown later below.

Pulse Duration

Pulse duration D of both a wide and a narrow pulse was measured. Figure 2 shows typical

OTDR traces from a high reflection (low ORL) fiber far-end. For low reflections, the

backscattering background (which increases for wider pulses) will distort and broaden the

reflected pulse, giving a less reliable value of pulse duration.

OTDR

dBs

W" IB

f 5 d

1

3

-W

.de I iilse

200m —W:

I3 ? irrov t Pill se

Ks

\

D{ns) = N W(m)/0.1499

TV = group index

W = OTDR pulse width

OTDR Distance (m)

Figure 2 - Wide and narrow OTDR pulses from a high reflection (low ORL)

Note that the pulses are not rectangular, and that there are few resolvable distance points

within a pulse. Therefore, the precise value of D is somewhat ambiguous. We measured pulse

duration at the usual [9] IV2 dB below peak power, corresponding to the FDHM of linear optical

power. In addition, since the backscatter power is proportional to pulse energy [8], we

measured D as the pulse energy (area) divided by the peak linear power. Results are given in

Table 1.

36
Wide Narrow

lYz dB duration

(FDHM) 2,127 377

energy/power

duration 2,499 489

Table 1 - Pulse durations D (in ns) for two definitions and pulse types

Pulse Height

A fiber reel with a reflecting far-end was fused to the OTDR pigtail, and the height P of

the far-end reflection peak above the backscatter level there was measured. For a highly

reflecting far-end, receiver clipping masked the true pulse height. Our solution to this was to

introduce a non-reflective bending attenuation into the fiber as shown by the signature in

Figure 3. This should not affect the true (undipped) height P.

OTDR dBs

OTDR Distance (km)

Figure 3 - A full-length OTDR trace, showing clipping and bending attenuation

However, for excessive attenuation, P decreased by up to 2 dB and 0.4 dB for 3Vz km and 5Ms

km fiber lengths, respectively. This is because the receiver's recovery tail due to the near-end

reflection [9] adds substantially (on the logarithmic scale) to the backscatter level, but adds

little to the end-reflection pulse peak. (The near-end recovery tail is not influenced by the

attenuation.) For the minimum (but still large) induced attenuation sufficient to prevent

clipping, height P was not noticably affected when the b

lk km fiber was used.

Fiber Backscatter

Three single-mode fiber types were investigated: depressed-clad, matched-clad, and one

having a large mode-field diameter (not typical of commercial production). Various reflectors

(i.e. cleaves) were characterised for each fiber by both coupler and OTDR methods. Then the

fiber backscattering level B was determined by adding the ORL from coupler Equation (2) to

the logarithmic term of OTDR Equation (3). Table 2 shows the B values for each fiber type.

37
ORL Backscatter B (dB)

Fiber (dB) Wide Pulse Narrow Pulse

TypeXT by coupler FDHM energy/power FDHM energy/power

matched- 14.4 79.3 80.0 78.9 80.0 * clad 25.7 80.2 80.9 79.7 80.9

35.1 80.5 81.2 79.6 80.7

depressed- 14.3 78.0 78.7 77.7 78.8 * clad 23.3 79.0 79.7 78.6 79.7

large mode 14.7 80.0 80.7 79.5 80.6

diameter 21.9 80.6 81.3 80.2 81.3 #

* average of 3 similar cleaves, # average of 2 similar cleaves

Table 2 - ORL and backcsatter values for several types of fiber and cleave

Note that for a particular fiber, the wide and narrow pulses give similar B values if the

duration is defined by pulse energy over pulse peak power. Values obtained with high

reflections (low ORLs) are slightly higher for reasons not fully understood. Note also that B

values can differ by up to 2 dB between fiber types.

Conclusions

We have shown that the OTDR can reliably measure return loss if care is taken in

obtaining reflected pulse parameters. Moreover, just as fiber group index needs to be known

for accurate distance measurements, it has been demonstrated that fiber backscatter level

needs to be known for accurate return loss measurements.

For purposes of ORL determination, we offer these suggestions for consideration by OTDR

manufacturers:

• Specify the pulse duration by the ratio of pulse energy to pulse peak power, rather than

by only the FDHM.

• Minimize the recovery tails of reflection pulses.

• Introduce an anti-clipping capability (e.g. variable attenuation) for reflected pulses.

Acknowledgements

The authors wish to express appreciation for valuable discussions with William W. Wood,

G. Shinn, and William C. Young.

References

[1] S.Wang, D.Burpee, & L.Jou, submitted to Electron. Lett.

[2] M.Choy, J.Gimlett, R.Welter, L.Kazovsky, & N.Cheung, Electron. Lett., v. 23, p. 1151 (1987)

[3] Bellcore Technical Reference TR-TSY-000326 (1987)

[4] EIA FOTP-107

[5] A.Judy & H.Neysmith, Fib. & Int. Opt, v. 7, p. 17 (1988)

[6] V.So, J.Lamont, & P.Vella, IEEE J. Sel. Area Comm., v. SAC-4, p. 737 (1986)

[7] This improves upon a result in [5].

[8] A.Hartog & M.Gold, J. Lightwave Tech., v. LT-2, p. 76 (1984)

[9] G.Shinn, Proc. EFOC/LAN 87, p. 151 (1987)

38
RETURN LOSS STUDIES OF A SINGLE-ENDED SINGLE-MODE FIBER:

EFFECT OF THE NONFLAT FIBER ENDFACE

V. Shah, W. C. Young, and L. Curtis

Bellcore

331 Newman Springs Road

Red Bank, NJ 07701-7020

(201)-758-3156

INTRODUCTION :

In optical fiber systems, optical feedback from reflections caused by index discontinuities such as

connectors and splices can generate intensity noise in lasers,'

1

' or interferometrically convert the intrinsic

phase noise of lasers into high levels of intensity noise.'2

' Both of these effects can adversely affect the

performance of direct detection and/or coherent lightwave systems.

When an index matching oil is used, improvement in the return loss value is limited due to the presence

of the polishing induced high-index region at the fiber endface.'3

' However, it is theoretically known'4'

that by tilting the fiber endface at an angle a with respect to a plane perpendicular to the fiber axis, the

amount of reflected power coupled to the fundamental mode of the single-mode fiber can be reduced by

several orders of magnitude (for a as low as five degrees). This results in very high return loss and

improves the system performance. Such a technique was recently used'5

' in multiple single-mode fiber

splices to achieve high return losses without adversely affecting the insertion loss performance. During

the course of that investigation it was observed that return losses of a single-ended single-mode fiber,

measured as a function of the tilt angle a, increased much less rapidly than predicted by the theory,'4'

particularly for values of a > 5 degrees. In this range of a, it was occasionally observed that, instead of

increasing, return loss values either did not change or in fact decreased when a larger tilt angle was

polished. These observations indicate that even though the return loss performance improves by

increasing a above 5 degrees, the improvement is not as significantly large as predicted'4'

The theoretical values in the previous work were obtained by treating the return loss in terms of the

coupling loss between two angularly misaligned gaussian beams assuming flat fiber endfaces. In this paper,

we show that experimental results discussed above can be explained by extending the theoretical model to account for the polishing-induced departure from flatness of the fiber endface. The results show that even

a small departure from flatness can have a large impact on the return loss performance. In addition, it shows the importance of optimum polishing and of proper wetting of the endface with index matching oil

in obtaining good return loss performance.

EXPERIMENT :

Return loss measurements were made with a 1540 nm wavelength laser and a fused biconical coupler. A

short ( ~ 1 m) length of fiber was fusion spliced at one output end of the coupler while the other output

end was terminated by polishing a 14 degree tilt angle on the endface and immersing it in an index

39
matching oil. Because of the presence of the polishing induced high-index region at the fiber endfacej3' an

oil with n = 1.6 was used Fig. 1 shows the measured values (o) of the return loss plotted as a function of

the tilt angle on the endface of the 1 m length of fiber. The solid curve represents the theoretical values

computed using Gaussian theoryJ

4

' It can be seen from Fig. 1 that the measured values differ considerably

from the theoretical values both for small and large tilt angles. We note that the solid curve was obtained

by assuming a perfectly smooth and flat endface. However, due to the differences between the hardness of

the core and the cladding material, polishing of the endface usually results in a depression in the core

region as shown in Fig. 2a. The high-index region present at the endface is modeled as shown by the

dashed line in Fig. 2a which represents the interface between the bulk material and high index regioa

The refractive index of this region is assumed to be uniform. When index matching oil is used to match

the high index region, only the interface between the bulk material and high index layer remains as shown in Fig. 2b.

THEORY :

In deriving the expression for power coupled in the fundamental mode due to reflections, we have used

the model in Fig. 2b. In addition to assumptions made above, we assume that : (1) The fundamental mode

has a Gaussian distribution with mode field diameter 2w, (2) That a is assumed to be small and, therefore,

simple Fresnel formula for normal incidence is used. Thus, with nc and ng denoting refractive indices of

the core and the oil, respectively, the power reflection coefficient RQ = [(n

g - nc )/(ng + nc )f, (3) The

depth of depression dQ

is small compared to the core diameter 2a, (4) The depression is assumed to be

spherical. The radius R of the sphere depends on dQ and is given by: R = (a 2 +dQ

2 )/(2dQ), and (5) The

cladding portion of the endface is assumed to be flat. With these assumptions, the field coupling

coefficient CQ for the fundamental mode is given by:

e

-i2kncg (a)j /{r)ei2kncg (r) ^ _

j /(r)

^

+ exp( _{

^cVvtana}2/2)] (1)

0 0 4

J

where / (r) = rexp(-2r 2/w 2

) J0(lknc ra\ and g(x) = (R 2 - x

2

) m. The return loss RL (in dB) is

obtained as

RL = -10 log

( |c 0

|

2 ) (2)

Fig. 3 shows the plot of RL as a function of a, with the solid curve from Fig. 1 repeated as curve A.

Curves B and C corresponding to do = 0.05 and 0.1 micron, respectively, show the effect of the very

small departure from the flatness of the fiber endface and were obtained by using Eqs. (1) and (2). The

crosses (x) represent the experimental values. Even though the simple model of Fig. 2b does not

correspond to the measurement conditions in all details because of the various assumptions discussed

above, good agreement between the measured values and our model is obtained when one notes that do

may have different values each time the endface is polished. The essential features of the experimental

observations are also closely predicted by our model which is therefore expected to play an important role

40
in providing valuable insights in optimizing high performance joints.

CONCLUSION:

In conclusion, even a very small polishing-induced departure from the flatness of the fiber endface results

in a much less rapid increase in return loss with increasing endface tilt angle compared to that obtained

with the flat fiber endface. The values of return losses as a function of endface tilt angle are in good

agreement with a nonflat endface model, thus indicating that the polishing-induced nonflatness of the

fiber endface plays a significant role in obtaining optimum return loss performance by tilting the fiber

endface.

REFERENCES:

[1] O. Hirota, Y. Suematsu, and K. S. Kwok, IEEE J. Quanttum Electron, QE-17, 1014 (1980).

[2] M. M. Choy, J. L. Gimlett, R. Welter, L. G. Kazovsky, and N. K. Cheung, Electron Lett 23, 1151

(1987).

[3] V. Shah, W. C. Young, and L. Curtis, OFC/IOOC87, TUF4, (1987)

[4] D. Marcuse, Bell Syst Tech. J. Vol. 56, 703 (1977)

[5] W. C. Young, V. S. Shah, L. Curtis, submitted to Electon Lett

CO

33,

CO

(/)

o

LU

CC 20

10

^^^^ O MEASURED — CALCULATED

n

1 2 3 4 5 6 7 8

ENDFACE TILT ANGLE (deg)

Fig. 1 Return loss for a single-ended single-mode fiber plotted as a function of the tilt angle. The endface

is immersed in an oil with n = 1.6. Calculated values are obtained by assuming the endface to be fiat

41
Fig. 2 (a) Fiber endface model with the high-index region at the endface. (b) Fiber endface model with

an index matching oil used to match the high-index regioa

DEGREES

Fig. 3 Return loss plotted as a function of the tilt angle. Curve A is the solid curve of Fig. 1. Curves B

and C correspond to do = 0.05 and 0.1 micron, respectively, and are obtained by using Eqs. (1) and (2) in

the text.

42
Optical Fiber Up-Taper Backreflector for Laser Transmitter Evaluation

Alan F. Benner, Daniel A. Fishman*, Herman M. Presby

AT&T Bell Laboratories

Crawford Hill Laboratory

Holmdel, New Jersey 07733

*AT&T Bell Laboratories

Holmdel, New Jersey 07733

INTRODUCTION

As optical communications systems progress to higher data rates, reflection

effects become increasingly detrimental to system performance. Reflections cause two

types of system degradation. First, optical feedback into the laser cavity causes

amplitude and phase fluctuations, introducing errors in the transmitted bit stream.

Second, when more than one reflection point exists in a system, doubly reflected light

interferes at the receiver with unreflected light, causing interference between different

time slots in the bit stream. Thus multi-path interference (MPI) has been shown to

cause data-dependent error bursts and bit-error-rate floors and is unaffected by laser

isolation.

In order to directly evaluate laser feedback effects, we have constructed a

variable back-reflector capable of producing uni-directional optical feedback in the

range -16.9 to -50 dB, without introducing observable MPI effects. The device utilizes

optical fiber up-tapers (Fig. 1) to reduce reflection levels in the apparatus to below

-50 dB, and an in-line optical isolator to eliminate spurious system reflections. We have

made measurements using this apparatus on optical feedback effects for various multilongitudinal-mode laser chips operating at 1.7 Gb/s.

EXPERIMENTAL ARRANGEMENT

The experimental apparatus is shown in Fig. 2. Elimination of secondary

reflections was made possible by the use of an in-line optical isolator and optical fiber

43
up-tapers with beveled end-faces'2

'.

Optical fiber up-tapers (Fig. 1) achieve single-mode beam expansion by a

gradual increase in the fiber cross-section, allowing adiabatic conversion of the singlemode fiber power into the (larger) fundamental mode of the expanded cylindrical

section. Up-tapers can expand the standard fiber mode diameter of 9 up to 50 or

100 /tm. The up-tapers used in this experiment had 55 //m mode diameters. Beam

expansion results in greater collimation of the emitted beam, relative to SM fiber, and

allows insertion of components between the tapers with minimal additional insertion

loss (<3 dB in this case). In addition, the common technique of beveling a fiber

termination to increase the return loss has a much greater effect for up-tapers than for

SM fibers, due to the expanded mode size'

3

', and the 3° bevel angle used here yielded a

return loss >50 dB.

The laser chip is coupled to the input up-taper using a microlens on the

fiber pigtail. A variable beamsplitter and variable attenuator produce the optical

feedback into the laser chip in the range -16.8 to -50 dB relative to the laser facet power

level. The mirror is mounted on a piezoelectric translator allowing 30 /zm travel, to

adjust reflected phase. A multi-mode fiber at the fourth port of the beamsplitter leads

to a PIN photodiode, allowing real-time measurement of the relative feedback power

and RF spectrum. Light collected at the receiving up-taper is passed through the

isolator and then through either a 6m length of 5D fiber jumper or through 3.1 km of

dispersion-shifted fiber before reaching a PIN diode direct-detection receiver. A

microcomputer interfaced with a BER test set is used to measure laser transmitter

performance under various reflection levels.

44
RESULTS

A sample BER measurement, made for a 1.3 micron multi-mode laser is

shown in Fig. 3. BER curves with and without 3.1 km of dispersion-shifted fiber, with

and without -16.8 dB feedback are shown. Intentional misalignment of up-tapers

showed no modal noise effects. With feedback and no dispersion, the laser transmitter

shows ~1.5 dB power penalty, with little excess curvature. With dispersion added, the

BER curve is drastically degraded, showing an error floor at 10

-8 BER. Varying the

reflected phase had no effect on BER results.

CONCLUSION

We have constructed a variable back-reflector which allows measurements of

the effects of optical feedback on a semiconductor laser, without introducing multi-path

interference effects. Optical fiber up-tapers greatly simplify alignment of the

apparatus, and allow very low reflection levels to be achieved more easily than with

standard SM fiber. This apparatus is the first use of up-tapers in transmission

experiments, and demonstrates the capability of up-tapers for reducing reflections in

optical fiber systems.

REFERENCES

1. J. L. Gimblett, J. Young, R. E. Spicer, N. K. Cheung, "Degradations in Gbit/s

DFB Laser Transmission Systems Due to Phase-to-intensity Noise Conversion by

Multiple Reflection Points", Electron. Lett., 1988, 84, pp. 406-408.

2. H. M. Presby, N. Amitay, A. Benner "Straight-Tip Optical Fiber Up-Tapers for

Single-Mode Hardware Applications", Electron. Lett., 1988, 84, pp. 34-35.

3. A. Benner, H. M. Presby, N. Amitay, "Low-Reflectivity In-Line Variable

Attenuator Utilizing Optical Fiber Up-Tapers", submitted to J. Lightwave

Technology.

45
SINGLE MOOE

FIBER PIGTAIL

Mirror

Variable urn

|

Attenuator

TAPERING

SECTION

CYLINDRICAL

SECTION

SPOT SIZE:

SPOT SIZE : 2u>o~ 50-100 fim

FIGURE 1 OPTICAL FIBER TAPERS

Up-Taper

w/ microlensed

pigtail

PD

Variable

Beamsplitter

Up-Taper

Monitor

Fiber

(multimode)

Ammeter

Optical

Isolator

5D Fiber - 6 m y\

I

RF

Spectrum

Analyzer

PD LJ DS Fiber • 3.1 km

Bit Error Rate

Test Set

Fig. 1 Experimental Arrangement

0-OX Rail. Oka daf

O-OX Raf 1, 3. lka daf

-S-2. IS Raf1, Oka daf

-A-2. IX Rafl. 3. lka daf

IE

IE

IE

IE

j r

Fig.

-37 -36 -35 -34 -33 -32 -31 -30 -Zc

. Received Optical Fou;er 'ciBm)

3 Multi-mode 1.3 um laser,

with and without dispersion

and reflection

46
Measurements of Polarization Mode Dispersion in Optical Fibers

Craig D. Poole

AT&T Bell Laboratories

Crawford Hill Laboratory

Holmdel, New Jersey 07733

The need for accurate measurements of polarization mode dispersion (PMD) in

single-mode fibers has been enhanced in recent years by the emergence of ultra-high bit rate

lightwave systems. To date, the vast majority of PMD measurements have been made in short

fibers. While such measurements have provided important information on intrinsic fiber

properties, it has been known for some time that in long fibers (i.e. > 100 meters ) extrinsic

factors, such as external perturbations created in spooling or cabling, can have a profound

impact on the dispersion through random mode coupling [1]. Until recently, the scarcity of

data in long fibers could be largely attributed to the lack of a practical model. The purpose of

this paper is to review recent measurements of PMD in long fibers using a recently developed

phenomenological model. With this model the effects of PMD, both in the frequency domain

and in the time domain, can be characterized in fibers of arbitrary length.

Historically, modal dispersion in multimode waveguides with random perturbations

has been treated using the coupled-power model in which the temporal response in optical

power of a waveguide is determined by considering the average distribution of power among

modes. For PMD this approach is impractical because source coherence times are often

comparable to or larger than the differential propagation delay times involved. Under these

conditions the transmission properties of a fiber become very sensitive to the source

frequency and to environmental conditions ( e.g. ambient temperature) through coherent

47
interference in the mode coupling process. Thus the coupled-power model, which neglects the

phase of the mode fields, can only predict the average behavior of an ensemble of fibers ( or

equivalently, the behavior of an individual fiber averaged over source wavelength and

environmental conditions).

Recently, a model of PMD for highly coherent sources was proposed in which the

dispersion in a fiber is described in terms of a differential delay time At associated with two

orthogonal input principal states of polarization [2],[3]. This description comes about as the

first order term in a frequency expansion of the transmission matrix of a fiber about the

source center frequency and, in the case of significant mode coupling, is restricted to the

regime of source spectral width Sv given by | Ar\6u « 1. In the absence of mode coupling,

the principal states are aligned with the polarization modes of the fiber and the differential

delay time is simply the differential group delay time for the modes. In the presence of

random mode coupling, however, the principal states and differential delay time are no longer

correlated with local fiber properties and can not be known a priori since they depend on the

details of the mode coupling. Furthermore, both become sensitive to the source wavelength

and to the fiber environment.

Since the principal states model can be viewed as a generalization of the eigenmode

picture in short fibers, many of the same measurement techniques that have been used to

measure modal group delay times in short fibers can be applied to the measurement of Ar in

long fibers. One approach that has been used in several recent experiments is to make

measurements in the frequency domain [4]-[7]. In the frequency domain, PMD manifests itself

as a frequency or wavelength dependence in the state of polarization at the output of a fiber.

This takes the form of a frequency dependent rotation of the output polarization on the

Poincare sphere, where the angular rate of rotation is exactly equal to the differential delay

time At. Figure 1 shows the rotation of the polarization at me output of a 147 km undersea

48
cable for three different launch conditions [7]. Note that the direction and magnitude of the

rotation is independent of the input polarization. The data is obtained by temperature tuning

the output of a DFB laser over an 1 8 GHz interval. While the frequency is being swept, the

state of polarization at the output of the cable is measured in real time using an electro-optic

polarimeter. Figure 2 shows the differential delay time between the principal states as a

function of wavelength for the four fibers in the cable. The curves are generated by

determining the rate of rotation of the polarization at each wavelength. Significant variation

© of the delay time over the 10 A interval can be seen in the upper curves. The estimated

accuracy in this particular experiment is limited by the DFB laser temperature monitor to ~ ±

10%.

The most obvious and direct approach to measuring PMD is to make time domain

measurements. Figure 3 shows the results of recent measurements of pulse propagation in 10

km of dispersion-shifted fiber [3]. Three output pulses are shown corresponding to the input

pulses having polarizations aligned with the two principal states (solid curves labeled i± ) and

with the input split equally between these states (dashed curve). A time shift of ~ 40 psec

between the the principal states is clearly seen. The intermediate waveform shows the

expected broadening of ~ 4% (indicated by a slightly lower peak). In this experiment, there is

clear evidence that mode coupling plays a significant role in the dispersion. This is indicated

by a slow drift in both the principal states and the differential delay time. Figure 4 shows the

change in the differential delay time over a twenty-four hour period. More than a factor of

two change is observed and can be attributed to small changes in ambient temperature.

As the data in Figure 4 illustrates, PMD in the case of highly coherent sources

ultimately requires a statistical treatment. For example, for a population of statistically

equivalent fiber, the rms value of the differential delay time is given by [8]

aAr = V<Ai2> = -4^-(e^2hL - 1 + 2/zL)1/2 where A/8' =-j-(J3x -ft.) is the differential

V2h du>

49
group delay time between polarization modes, L is the fiber length, and h is the well known

h-parameter that characterizes the coupling between modes due to random perturbations. The

symbols < > indicate an average over the ensemble of fibers. At short lengths the dispersion

reduces to the differential group delay time between the modes, limtrAT = A/3'L, and at long

KL—O

lengths exhibits the expected square root of length dependence, lim aAr = A/3'VL /h . hL—*oo

Random mode coupling therefore has the effect in long fibers of reducing the dispersion by a

factor of VhL relative to the no mode coupling case. It has also been shown that the

probability density function for Ar approaches a Gaussian shape at long lengths [8].

Under the usual assumption of ergodicity, measurement of crAr or alternatively


 can be accomplished by averaging measurements over some physical parameter such

as temperature or source wavelength. In other words one assumes that

<| Ar|>em, emWe = 
 r = 
A. Figure 5 shows the apparatus used recently to measure

the length dependence of 
 in conventional single-mode fiber [9]. Measurements are

made by scanning the wavelength of the source between 1.2 and 1.6 /xm while monitoring the

power transmitted through an analyzer positioned at the output of the test fiber. The

oscillations in the transmitted power reflect the rotation of the polarization and are used to

determine the dispersion over a broad wavelength interval. Figure 6 shows the results of cutback measurements made on two pieces of fiber taken from a single spool. To determine the

effect of external perturbations, one of the pieces is wound on a conventional 1 5 cm diameter

spool ( open circles ) while the other is wound at low tension in a single layer on a 28 cm

diameter drum (solid points). Each data point represents the average of |Ar| over the

wavelength interval 1.2 -1.6 /zm. A theoretical fit is performed for each set of data using the

values of h-parameter shown and the measured modal group delay time of 2.8 psec/km. The

importance of external perturbations is demonstrated by the more than order of magnitude

difference between the data on the drum and that on the spool ( two orders of magnitude in

50
the h parameter ). The transition from linear to square root dependence in the length is

clearly seen in the drum data while the spool data exhibits a square root dependence down to

the shortest length due to the much greater mode coupling.

In conclusion, PMD in long fibers is sensitive to such variables as source spectrum,

mechanical perturbations on the fiber, and environmental conditions (e.g. ambient

temperature). All of these factors need to be considered when making measurements. A

statistical treatment of the dispersion based on the ensemble average of the differential delay

time between principal states is one approach to obtaining a meaningful characterization of

the dispersion that is both measurable and relevant to lightwave system penalties.

REFERENCES

[1] K. Mochizuki, Y. Namihira, and H. Wakabayashi, Electron. Lett. 17, 153 (1981)

[2] C. D. Poole and R. E. Wagner, Electron. Lett., 22, 1029 (1986).

[3] C. D. Poole and C. R. Giles, Opt. Lett. 13, 155 (1988)

[4] N. S. Bergano, C. D. Poole, R. E. Wagner, IEEE J. Lightwave Technol. LT-5, 1618

(1987).

[5] D. Andresciani, F. Curti, F. Matera, and B. Daino, Opt. Lett. 12, 844 (1987)

[6] M. Tsubokawa and Y. Sasaki, Electron. Lett. 24, 350 (1988)

[7] C. D. Poole, N. S. Bergano, R. E. Wagner, and H. J. Schulte, IEEE J. Lightwave Technol.

LT-6, 1184 (1988)

[8] C. D. Poole, Optics Lett. 13, August Issue (1988)

[9] C. D. Poole, Submitted to Optics Lett.

51
12

Ot 1 1 I I I I 0 2 4 6 8 10 12 Fig. 1. Frequency dependent polarization change in wavelength <a> out of 147 km cable.

Fig. 2. Differential delay time for four

fibers in 147 km cable.

CDE

<

a

55

_iS

i

o

_ATT

1 r i i i

DISPERSION-SHIFTED

RBER

L = 10 km

X = 1.55

A-*| \*~ HO f <~.

INTERMEDIATE

ft/ V\

ffv \VV\ *

V— ZERO

i i i i i i i i i TIME (100 ps/dlv)

Fig. 3. Pulses transmitted through 10 km

of dispersion-shifted fiber.

1 T 1 FIBER 1

T 1 1 1 1 1

" L = 10 km —»j U 22 p«ec

^^^^^^

i

i

i

i

i

-H |«- - 55 p»ec

f

24 HRS y A\

LATER

1 1 1

i i i i i i TIME (100 ps/dlv)

Fig. 4. Drift of differential delay

time over 24 hours.

1 f=

LAMP

CHOPPER

MONOCHROMATOR 0

FIBER

POLARIZER

COMPUTER LOCK-IN

|—<T| Kh[

Ge ~

DETECTOR

Fig. 5. Set-up for measuring polarization

dispersion.

• 28 cm DIA. DRUM /LINEAR

)

o

<D

01Q. <z> A 0.1 K

S

<\V

0.01

o 15 cm DIA. SPOOL /

THEORETICAL /

100 1000

LENGTH (METERS)

10,000

Fig. 6. Length dependence of dispersion.

52
HETERODYNE FIBER POLARIMETRY WITH STOKES PARAMETER DETECTION

R. Calvani, R. Caponi, F. Cisternino

CSELT - Via Reiss Romoli 274 - 10148 Torino - Italy

Abstract

A novel heterodyne polarimeter for Stokes parameter measurement is

demonstrated. A two parameter output, tested with a known polarization phase

input, gives expected results. A precise beat-length measurement on a high

birefringence fiber is performed.

Introduct ion

Heterodyne polar imetry is a recent measurement method particularly suitable for

a fast detection of the state of polarization (SOP) of the light at the output

of an optical fiber [1]. The inherent principle is the same as in coherent

transmissions, where the optical information carrier is scaled into the

radio-frequency (RF) range through a beating with a radiation at a slightly

different frequency (local oscillator). This techique, applied to both

components E x , Ey of an optical coherent field, provides two electrical RF

signals which reproduce phase (J> and amplitude ratio R of the SOP to be

measured. In a more general case, when depolarization phenomena occur (for

example in polarization crosstalk measurements on long fiber spans), E x and Ey

are no more sufficient to identify the SOP, which requires the four independent

Stokes parameters Sq, s-| , S2, S3.

To this end, the proposed heterodyne polarimeter has been developed to provide

directly at the output (with analog electrical processing) the aforesaid

parameters. In the present paper, first experimental results of this novel

instrument version concerning S3 are discussed. A test with an external

stress apparatus on a polarization maintaining fibre has given highly precise

measurements of the submi 1 1 imeter beat-length.

The Stokes polarimeter, being a device suitable for a complete characterization

53
of the statical and dynamical polarization behaviour of single-mode fibers, can

be usefully applied in advanced research areas, such as coherent communications

and optical sensors [2].

Measurement principle and exper imental implementat ion

In heterodyne polarimetry, two RF signals are obtained from the beating of a

reference field (linearly polarized at 45°) with the two measurand E x , Ey [1]

(1) S x

oc E 0x cos (fit) S

y

x E 0y cos (Qt-cM

where </> is the polarization phase. In the present system, the reference beam

is derived from the same source, in a conventional interf erometr ic setup

(Mach-Zehender configuration), where an acousto-opt ic device introduces the

frequency shift Q (Fig. 1).

At this point, simply by mixing the electrical signals S x , Sy (and filtering

the 2Q component), one obtains a quantity proportional to S2 parameter; S3 can

be derived the same way, delaying either of the two RF ' s by 90°

(2) s 2 = 2 E 0x E 0y cos (f> s 3 = 2 E 0x E 0y sin <f>

For the s parameter a different processing of S x , Sy is required; the two RF's

are separately square low detected and finally subtracted

(3) Si = E 0x

2 _ E 0y

2

This kind of signal treatement has already found interesting applications in

the coherent communication systems which use polarization modulation as

information coding [3,4].

On the other hand, the s energy parameter has to be obtained by a radiometric

measurement. The polarized energy ( S-] 2 + s2^+S3^ ) is given by the sum of the two

square law detected outputs (Sx ^+Sy2 ) and its ratio with respect to the total

energy provides the polarization degree.

To test the effectiveness of the proposed method, the scheme of Fig. 1 has been

chosen. The source is a single frequency He-Ne laser; the fiber is

polarization maintaining (bow tie), with 513 nm cutoff wavelength, fabricated

54
by York Technologies. A stress applying device for fiber measurement is also

added

Results and d i scuss ion

A first test on this heterodyne polarimeter has been performed by using the two

output signals Sx , Sy in a conventional way, i.e. by sending the two RF

beatings to an x-y oscilloscope, in order to visualize directly a known SOP

ellipse (Fig. 2). This way is also possible to calibrate the instrument,

compensating for spurious birefringences of optical components [1].

The new electrical processing scheme, in terms of S2, S3 parameters, has been

checked using a known SOP phase modulation (180° depth), provided by a Pockels

cell placed on the measurement arm. The resulting curve on S2, S3 plane has

been visualized on the x-y oscilloscope (in this case the output signals

oscillate at the SOP modulation frequency, because the RF's are subtracted at

the mixers). Because of the 180° phase modulation, the resulting curve is an

half circumference, as espected (Fig. 3).

The fiber measurement (beat-length Lg ) has been performed by scanning a short

fiber span with the movable stress device, which heavily modifies the local

SOP. The output signal periodicity on either of S2, S3 parameters gives the Lg

value [1]. From the measured data (Fig. 4), the averaged Lg is 0.94 mm in

excellent agreement (1%) with the value measured at the factory. The

possibility of choosing between different Stokes parameters, provided by this

instrument, makes this measurement independent of the SOP at the fiber input

(so avoiding the search of the eigenaxes).

References

[1] R. Calvani, R. Caponi, F. Cisternino, G. Coppa, JLT, LT-5 , 1176 (1987).

[2] B. Culshaw, Optical Fiber Sensing and Signal Processing, London, 1984.

[3] R. Calvani, R. Caponi, F. Cisternino, El. Lett., to be published.

[4] E. Dietrich, B. Enning, R. Gross, H. Knupke, El. Lett., 2J3 , 421 ( 1987 ).

55
STRESS DEVICE

HETERODYNE STOKES POLARIMETER

Fig. 1 - Experimental setup for the Stokes pol ar imeter

f

Fig. 2 - x-y visualization of an output SOP ellipse

Fig. 3 - Display on s 2 , s 3 Stokes

parameter plane with a 180° polarization phase modulation

Fig. 4 - Experimental data of s 2

Stokes parameter versus axial

scanni ng

56
COMPLETE BIREFRINGENCE CHARACTERISATION OF LITHIUM

NIOBATE INTEGRATED OPTIC POLARISATION CONTROLLERS

G R Walker, N G Walker

British Telecom Research Laboratories

Martlesham Heath, Ipswich IP5 7RE, United Kingdom

Abstract: A computer- controlled system for the measurement of

birefringence is described. Results are presented from a study of a novel

lithium niobate polarisation controller.

Introduction: Many optical devices are polarisation dependent, and a means

of readily characterising their birefringence is needed. A specific

example is polarisation controllers, which are important components for

coherent optical transmission systems and certain optical fibre sensors.

Various lithium niobate waveguide polarisation controllers have been

proposed and fabricated, but assessment of their performance has typically

been limited to measurement of conversion efficiency by power measurements

along specific axes [1,2,3]. This may be sufficient for simple structures

and applications, but more detailed characterisation of device performance

is clearly desirable. This paper describes a novel approach which gives a

complete birefringence characterisation, and which has been successfully

applied to two lithium niobate device structures [4,5].

Experimental Arrangement: The system for measuring birefringence is shown

in figure 1. Linearly polarised light passes through a half -wave plate,

which rotates to allow any linear state of polarisation (SOP) to be

launched into the device being characterised. The resulting output SOP is

measured using a rotating quarter-wave plate polarimeter [6] . Represented

on the Poincare sphere, the device birefringence corresponds to a rotation

about the birefringence axis through an angle given by the magnitude of

birefringence [6] . Measurement of the output SOPs for at least two

different known input SOPs enables this birefringence to be calculated,

although four launch states are used in practice to give some averaging.

Since any rotation of magnitude 2ir radians leaves all SOPs on the Poincare

sphere unchanged, the magnitude of the birefringence can only be measured

modulo 2ir. However, by assuming that the birefringence is a continuous

function of the device control voltages, this ±2tt ambiguity can be

resolved, and the device birefringence is then completely characterised.

57
R0TATA8LE

HALF-WAVE PLATE D

FBRE-PIGTALED

DEVICE UNDER TEST

LINEARLY-POLARISED

SOURCE

Q

DEVICE

CONTROL

VOLTAGES

ROTATING

QUARTER -WAVE

PLATE POLARI METER

LAUNCH SOP

CONTROL

4—

Figure 1 Measurement system

COMPUTER

POLARI METER

CONTROL

SOP MEASUREMENTS

4

The technique can also be used if there is a fixed birefringence

between the generated launch states and the device to be characterised,

and/or between the device and the polarimeter, as would be the case for a

fibre pigtailed device, or an electrooptic device with sections of

waveguide extending beyond the electrode pattern. It is possible to

compensate for these unwanted rotations by referencing the measured output

states to the output SOPs observed for the condition of zero birefringence

of the device under test. When using this approach, any error in

determining the zero of the birefringence of the device will cause large

errors in the calculated birefringence axis orientation for birefringence

angles close to 2n , hit and so on. This problem is avoided by presenting

the data as birefringence vectors with length equal to the sine of half the

birefringence angle, an approach which may also be significant in assessing

the control capabilities of the device.

Results: A novel lithium niobate integrated optic polarisation controller

has been thoroughly characterised using the arrangement described above.

The operation of the device has been reported in detail elsewhere [4] . In

outline it is expected to give a birefringence vector of variable magnitude

which can be rotated in the equatorial plane of the Poincare sphere. A

scan of 1,600 birefringence measurements takes about an hour, and the

results obtained for a typical waveguide are shown in figure 2(a). The

four surfaces are the three components of the birefringence vectors with

length equal to the sine of half the birefringence angle, and the cosine of

half the birefringence angle, plotted as functions of the two electrode

control voltages. Figure 2(b) shows the corresponding results from a

computer simulation of an "ideal" device. There is remarkable agreement

58
(a)

(b)

Figure 2 Birefringence characterisation of a lithium niobate polarisation

controller: (a) experimental, and (b) computer simulation of an

"ideal" device. Plots (i), (ii) and (iii) are the components of

the birefringence vector with length equal to the sine of half

the birefringence angle, and plot (iv) is the cosine of half the

birefringence angle.

59
over the full birefringence range of over 3tt . As an example of the detail

which can be revealed by this technique, the deviation of the Z- component

from zero was identified by further computer simulations to be caused by a

slow variation in the intrinsic waveguide birefringence along the length of

the device. It is interesting to note that the bulk-optic fractional

wave -plates in the measurement system could potentially be replaced by this

lithium niobate device, greatly increasing the speed of operation of the

birefringence analyser.

Conclusions: A novel technique has been described for complete

birefringence characterisation of optical devices, and its application to a

waveguide polarisation controller has been demonstrated. It has been

applied to two different electrode patterns on lithium niobate, and has

produced results in excellent agreement with the predicted performance.

Such detailed assessments of device performance can be obtained for a wide

range of optical devices, and will become increasingly important for

complex polarisation controller and filter structures.

Acknowledgements: The authors thank Jonathan Davidson and other colleagues

in the Lithium Niobate Device Development Group for fabricating the devices

characterised. The Director of Research and Technology, British Telecom

Research Laboratories is acknowledged for permission to publish this paper.

References

1 Alferness R C: ' Electrooptic Guided-Wave Device for General Polarization

Transformations', IEEE J. Quantum Electron, OE-17 . No 6, pp965-969,

(1981) .

2 Heidrich H, Helmholt C H, Hoffman D and Ahlers H: 'Integrated optical

compensator on Ti:LiNb03 for continuous and reset- free polarization

control', Proc. ECOC , Finlandia Hall, Finland, Tech. Digest 1,

pp257-260, September 13-17, (1987).

3 Donaldson A and Wong K K: ' Phase -matched mode convertor in LiNb03 using

near-z-axis propagation', Electron Lett., 23, No 25, ppl378-1379,

(1987) . 4 Walker N G, Walker G R, Davidson J, Cunningham D C, Beaumont A R and

Booth R C: 'Lithium Niobate Waveguide Polarisation Convertor', ibid, Ik,

No 2, ppl03-105 (1988)

5 Walker N G, Walker G R and Booth R C: 'Performance of lithium niobate

polarisation convertor/tunable filter', ibid, 24, No 5, pp268-270,

(1988) .

6 Walker N G and Walker G R: 'Polarisation Control for coherent optical

fibre systems', Br. Telecom Technol. J., 5, No 2, pp63-76, (1987).

60
TECHNIQUES FOR REFRACTIVE INDEX MEASUREMENTS ON

SEMICONDUCTOR WAVEGUIDES: PRISM COUPLING AND MODAL

CUT-OFF SPECTROSCOPY

C. De Bernardi, M. Meliga, S. Morasca, B. Sordo

CSELT S.p.A. - Via G. Reiss Romoli 274 - 10148 Torino, Italy

Introduction

Integrated optical waveguides on semiconductor materials are in rapid development, for the

great potential offerd by integrated optoelectronics, both in optical communications and in

optical data processing. Waveguide attenuation, in both GaAs and InP based materials, has

dropped well below 1 dB/cm, and is practically equal to the values exhibited by lithium

niobate and other waveguide materials: this makes it practical to consider integration of

several components and devices, requiring relatively long paths, on a single chip, without

incurring in intolerable losses or requiring on-chip amplification.

The design of sophisticated components needs accurate knowledge of the refractive indices

both of the substrate and of the layers involved in the propagation of radiation (waveguide,

cladding and cover layers etc.), which depend not only on the composition of the material,

but also on the doping levels; unfortunately, so far the available data on this aspect are not

as rich as needed, from the point of view both of quantity and precision, and device design

relies more on estimates from theoretical models than on precise experimental

determinations of these physical parameters.

In this paper we present results based on two techniques developed to improve the

knowledge of these refractive indices: the first one is an extension of the well-known prism

coupling technique to high index materials, and uses a silicon prism to determine the

refractive index of substrate material or of epitaxial layers, with an accuracy to the fourth

decimal place. The second method, modal cut-off spectroscopy (MCS), yields the refractive

index of waveguides at the wavelengths where cut-off of a guided mode is detected, if the

substrate refractive index is known; the accuracy reaches the third decimal place.

Prism coupling measurements

The use of a prism to couple radiation into waveguides is well known [1], and is widely

applied to the measurement of modal indices in planar guides. However, so far this

technique has not been used, to our knowledge, in the case of semiconductors, probably due to

their high refractive index and brittleness. We have used a silicon prism, suitable for InP

and a range of InGaAlAs and InGaAsP compositions, due to its high index of refraction; in our

case the prism data are: refracting angle a=64.25°, refractive index at 1.31 |im n=3.5039 ,

calculated from a Herzberger interpolation formula [2] (direct measurement of this value by

61
minimum deviation is in progress); presently a diode laser source at 1.31 urn is used, but

extension to other wavelengths with a color center laser is planned.

To characterize InP substrates the critical angle between prism and wafer is measured,

observing from the cleaved edge the beam refracted into the wafer; data for samples with

different doping are reported in Table 1: the accuracy reached in our case on thin wafers is

similar to that provided by minimum deviation measurements on bulk material [3], without

requiring special fabrication, or sample preparation, or waste of costly material.

For planar waveguides, coupling angles were measured for InGaAlAs on InP: results are listed

in Table 2 for a guide 2.7 u.m thick, grown by molecular beam epitaxy. In this case the

synchronous angles appear rather broadly defined in comparison to glass or LiNbC>3 guides,

up to about 0.1° instead of the usual value of a few 0.001°: this aspect is being investigated.

Table 1. Refractive index (n)

of InP wafers at 131 urn

Sample Doping n

#1 S 3.1634 ± 0.0003

#2 Fe 32055 ± 0.0003

#3 Sn 32009 ± 0.0002

#4 undoped 32063 ± 0.0002

#5 Sn 32008 + 0.0003

Table 2. Effective modal indices for

an InGaAlAs/InP waveguide

Mode ngff

TEo 33274 ±0.0002

TEi 33083 ±0.001

TE2 32714 ±0.001

TE3 3.223 ±0.002

Modal cut-off spectroscopy measurements

The spectrum of modal cut-offs in planar waveguides is function of the refractive index

difference between substrate and guiding layer and of its thickness, for step-index guides as in

the present case; for graded-index ones, also on the profile shape. Since the substrate index

can be obtained by the previous prism technique and the guiding layer thickness can be

accurately measured, the determination of the cut-off spectrum in combination with

theoretical dispersion curves yields values of the refractive index of the guide layer. An

extension of the theoretical model to laterally confined waveguides can be used to obtain the

same optical parameters for stripe guides [4,5]. This method has the advantage of operating

directly on stripe guides, providing results accurate to the third decimal figure, at each

wavelength of detected cut-off, which appears as a peak in the insertion loss spectrum of the

guide. Measurements have been performed on InGaAsP and InGaAlAs ridge guides grown by

liquid phase epitaxy (LPE) and molecular beam epitaxy (MBE) respectively. A schematic of

the experimental set-up is shown in Fig. 1: a stabilized tungsten-halogen lamp is filtered by

a scanning monochromator, and the monochromatic light is coupled to a single-mode fibre,

which in turn is butt-coupled to the guide under measurement; the output from the guide is

focused onto a photodetector, whose current is measured by a lock-in amplifier. An example

of spectral scan is shown in Fig. 2, while numerical results are listed in Table 3.

62
Table 3. Refractive indices of epitaxial layers obtained by MCS

Material Thickness A,gapum cut-off X TE An^ cut-off^ TM An^ n

MBEInGaAlAs 3.2 um 1.04 1.14 um 0.872 1.12 urn 0.883 3.395

1.44 urn 0538 1.40 um 0.817 3314

LPE InGaAsP 3.6 um 1.10 1.31 um 0.970 1.30 um 0.976 3.351

1.61 um 0.965 1.58 um 0.967 3310

In this table An^ is the difference between the refractive indices squared of the epitaxial

layer and of the substrate, and is directly obtained from the measurement, as illustrated in

Fig. 3; the refractive index n is derived from the relevant InP value calculated by an

interpolation formula forced through the value at 1.31 um. The uncertainty in n provided by

this method depends on the uncertainty of the epilayer thickness and of the cut-off

wavelength; typical values of 50 nm and 10 nm respectively lead to an accuracy of +.0.002 .

Discussion

The two techniques described here are complementary; the accuracy of their results, besides

allowing more sophisticated design, is also helpful to determine the composition of

compound materials, e.g. by using a modified single oscillator model, which connects

refractive index to material composition [6]. Prism coupling requires monochromatic, high

intensity sources and offers very high precision, comparable to bulk techniques, on standard

wafers, while MCS can yield results at any desired wavelength, by a proper choice of guide

thickness, which affects the cut-off wavelengths. Both techniques are insensitive, in

comparison to ellipsometry and reflectometry, to surface oxide films.

Conclusions

Results are presented for two refractive index determination techniques applied to

semiconductors, namely silicon prism coupling and modal cut-off spectroscopy; acccuracy to

the fourth decimal place is demonstrated, as well as operation directly on stripe

waveguides. This allows more accurate design of elaborate integrated optoelectronic devices.

Acknowledgments

The authors thank G. Fornuto, C. Rigo and A. Stano for providing the semiconductor samples.

This work has been partially supported by EEC under Esprit Project 263.

References

[1] T. Tamir, "Integrated Optics", (Springer Verlag, Berlin, 1975), p. 22.

[2] L. Levi, "Applied Optics", (J. Wiley & Sons, New York, 1980), Vol. 2, p. 911.

63
[3] E. D. Palik, "Handbook of optical constants of solids", (Academic Press, Orlando, 1985),

p. 504.

[4]M. Meliga, S. Morasca, B. Sordo, C. De Bernardi, "Spectral cut-off behaviour of Ti:LiNbC>3

and semiconductor waveguides: analysis and experiment", SPIE Proc. 835, p. 251-256, 1987.

[5] S. Morasca, B. Sordo, C. De Bernardi, M. Meliga, "Refractive indices of semiconductor

films measured by waveguide cut-off spectroscopy",Appl. Phys. Lett. 52, 1593-1595 (1988).

[6] B. Broberg, S. Lindgren, "Refractive index of Ini_x Gax Asy Pi-y layers and InP in the

transparent wavelength region", J.Appl.Phys. 55, 3376-3381 (1984).

M0N0CHR0MAT0R

LAMP GUIDE BEAMSPLITTER DETECT0R

FIG. 1. Schematic of the optical setup for modal cut-off sprctroscopy

•TM polarization

TE polarization

10

10

10

10

5 modes —

4 modes —

3 modes —

2 modes

1 mode

i i 1

i i i I i i—i—I—i—i—i J 1.2 1.4 1.6 1.8

A (urn)

Fig. 2. Measured cut-off spectra on a InGaAsP Fig. 3. Cut-off curves calculated for the guide

ridge guide 10 |im wide, 3.6 urn thick. of Fig. 2 for TM polarization.

64
ELECTRICALLY CALIBRATED PHOTOTHERMAL DEFLECTION MEASUREMENT

FOR SEPARATING CHANNEL WAVEGUIDE LOSS MECHANISMS*

Robert K. Hickernell, Donald R. Larson and Robert J. Phelan, Jr.

National Bureau of Standards, Electromagnetic Technology Division

Boulder, CO 80303

Introduction

Absorption and scattering are the two primary sources of attenuation in

optical waveguides. Their respective contributions to loss may differ

widely, depending on the wavelength, substrate material, fabrication method

and processing parameters. While the understanding of propagation loss is

important to the further reduction of insertion loss in waveguide devices,

it appears that little experimental research has been directed toward the

separation of loss mechanisms.

Some studies have analyzed the relative angular distribution of out-ofplane scattered light or the total waveguide attenuation using surface and

volume scattering theory [1,2]. However, the absolute measurement of power

scattered from a waveguide, both in-plane and out-of-plane , is very difficult. Absorptive loss in Ti:LiNb03 waveguides has been measured using

calorimetry [3] and the pyroelectric effect [4] , but published values of the

relative contributions of absorption and scattering disagree significantly.

Recently we applied the photothermal deflection (PTD) effect [5] to the

measurement of total propagation loss in channel waveguides [6]. The method

is noncontact and applicable to a wide range of channel waveguide geometries

and materials, including buried guides. Since it is based on the absorbed

power as a relative indicator of the guided power, it has an advantage over

methods based on scattered light in that scattering centers and unguided

background light do not affect the PTD measurement directly. With an

absolute calibration of absorbed power, it is possible to use the PTD method

to measure absorptive loss in waveguides. In this paper we present a

technique for distinguishing absorptive from scattering contributions to

channel waveguide loss by electrically calibrating the PTD effect. The

technique is illustrated by experimental results with ion-exchanged glass

waveguides

Contribution of the U.S. Government, not subject to copyright.

65
Measurement Apparatus and Techniques

The experimental apparatus used to make loss measurements in waveguides

is shown in Fig. 1, the details of which are described in Ref. 6. In

summary, the pump laser light is chopped and coupled into a waveguide, and

its absorption induces a modulated temperature and index gradient near the

guide. The gradient refracts a probe beam, at 633 nm wavelength, by an

amount measured by a bicell detector and lock- in amplifier. We present data

in the form AV/V, the ratio of the ac and dc signals of the bicell.

Measurements were performed on straight, potassium ion- exchanged, channel

waveguides in soda-lime glass substrates. The substrates were masked with

10-/zm channel widths, and the resulting guides were multimode at a wavelength of 633 nm.

Calibration of the PTD signal was accomplished by mimicking the

absorbed waveguide power with a thin film heater. We deposited NiCr stripes

160 nm thick, 5 /xm wide and 10 mm long on the glass waveguide substrates and

connected them to a square-wave signal generator, as shown in Fig. 2. The

resistance of a stripe was 63 kO. The probe-beam optics were the same as

those used to measure the PTD signal originating from absorbed, guided

light. When a known voltage was applied to the measured resistance, a

calibrated amount of heat was generated in a profile which approximated the

generation of heat by guided light.

Experimental Results and Discussion

For a probe spot diameter of 25 /zm and 633 nm guided light chopped at

500 Hz, the normalized PTD magnitude is plotted in Fig. 3 as a function of

the probe position transverse to a typical waveguide. The PTD signal was

zero at the center of the guide, where there was no thermal gradient. It

peaked at a magnitude of 3.8 x 10"^ at a position approximately 15 /xm from

the center of the guide. We monitored the peak PTD signal as a function of

position down the length of the guide and plotted the results in Fig. 4.

Linear regression indicates a total propagation loss of 1.2 ± 0.2 dB/cm.

The uncertainty is caused not by measurement error but by local absorption

inhomogeneities in the waveguide [6]

We measured the throughput power of the waveguide, accounted for

reflection losses , and calculated the guided power at the location along the

waveguide where the measurements of Fig. 3 were taken. The peak of the

66
chopped power was 2.9 mW; therefore, a peak of 0.81 mW/cm of light was lost

at that location due to both absortion and scattering.

In the electrical calibration procedure the voltage across the wire was

a positive square wave with a 4.0 V amplitude and a frequency of 500 Hz,

which generated a peak 0.26 mW/cm of heat. A transverse scan of the normalized PTD signal due to the wire heater is shown in Fig. 3 for comparison

with the absorption- induced signal. The shapes of the two scans are very

nearly the same, indicating that the wire heater approximately models the

waveguide absorption. The PTD signal from the thin- film heater was identical for different positions along the stripe and peaked at 6.4 x 10"^.

Waveguide absorption loss as a fraction of total loss is calculated as the

ratio of the PTD signals for the waveguide and the resistive heater,

normalized to their respective powers lost per length. The absorption

fraction measured here was 0.19 ± 0.05. This value is comparable to the

bulk propagation loss measured in the substrate, which was 0.12 times the

waveguide propagation loss. The ion-exchange process adds to the bulk loss

both scattering and absorptive contributions.

Uncertainty in the absolute calibration of absorption loss is due to

uncertainty in the measurement of total loss. As mentioned above, for the

waveguides used here the uncertainty is caused by the random distribution of

inhomogeneities in the waveguide material. Waveguides in higher quality

substrates are expected to exhibit a lower uncertainty. Studies of other

waveguides and at near- infrared wavelengths are currently being performed.

References

[1] M. Gottlieb, G.B. Brandt and J.J. Conroy, IEEE Trans. Circuits Syst.

CAS-26, 1029 (1979).

[2] G.H. Ames and D.G. Hall, IEEE J. Quantum Electron. QE-19, 845 (1983).

[3] S.D. Allen, E. Garmire, M. Bass and B. Packer, Appl. Phys . Lett. 34,

435 (1979)

[4] A.M. Glass, I. P. Kaminow, A. A. Ballman and D.H. Olson, Appl. Opt. 19,

276 (1980).

[5] W.B. Jackson, N.M. Amer, A.C. Boccara, and D. Fournier, Appl. Opt. 20,

1333 (1981)

[6] R.K. Hickernell, D.R. Larson, R.J. Phelan, Jr. and L.E. Larson, Appl.

Opt., July 1988 (to be published).

67
PUMP

LASER B-/%-0-Ti - v q

WAVEGUIDE

DETECTOR

BICELL DETECTOR -7^

DIFF. LOCK-IN

AMP. AMPLIFIER

SIGNAL GENERATOR

THIN FILM HEATER

BICELL DETECTOR

DIFF. AMP. LOCK-IN

AMPLIFIER

Fig. 1. Crossed-beam configuration

for photothermal deflection measurement of a waveguide.

Fig. 2. Apparatus for electrical

calibration of photothermal deflection.

1.00

<D "O

3

'c

o> 0. o

E

75

Q_ 0.50

X>

Q)N

| 0.25

i_O

c

0.00

+

-b

+

b

o

+

o

+40

o +

0 20 40 60 80

probe beam offset (Aim)

XI

cnD

E

0.50

0.25

0.00 -

CL

-0.25

O

-0.50

1—i—l—i—l—l—l—l—l—l—l—I—l—l—l—l—i—i—i—l—i—i—r—ra = 1 .2 ± 0.2 oB/cm

-i i i i i i i i i i i—i—i—i—i—i—i—i—i- 2.50 2.00 1.50 1.00 0.50 0.00 longitudinal position (cm)

Fig. 3. Normalized PTD signal

magnitude as a function of probe

beam offset from waveguide or wire

center: waveguide (+) ; wire heater

(o).

Fig. 4. Logarithm of the peak PTD

magnitude as a function of distance

along the waveguide. Best- fit line

yields a propagation loss of 1.2 ±

0.2 dB/cm.

68
Fiber Measurements for Long Term Reliability: Strength, Hydrogen, and Radiation

F. V. DiMarcello, P. J. Lemaire and J. R. Simpson

AT&T Bell Laboratories

Murray Hill, New Jersey 07974

INTRODUCTION

The long-term reliability of glass fiber lightguides acquires increased importance as system requirements

become more stringent. In addition, lightguides are finding applications in adverse environments which tend to accelerate many degradation mechanisms. Three important reliability issues(1) have emerged: the degradation of

fiber strength with time, optical loss increases due to the permeation of H2 into the fiber, and optical effects

resulting from the interaction of various types of radiation with the fiber. The design of meaningful experimental

methods as well as the development of accurate measurement techniques play important roles in predicting longterm reliability.

STRENGTH

At liquid nitrogen temperatures, where moisture is eliminated from the testing environment, silica fiber

strengths of approximately 2xl06 psi (13.8 GPa) have been measured.(2) At ambient temperature and moisture

conditions, however, this intrinsic value is reduced to approximately 800 ksi (5.5GPa) in short gauge lengths as a

result of the fatigue which takes place during the measurement. Furthermore, it is found that as longer lengths of

fiber are tested, the strength distribution is basically bimodal with a narrow high strength mode at approximately

800 ksi (5.5 GPa) and a broad low strength tail. This low strength mode is clearly extrinsic and is explained by

the classic Griffith model(3) of brittle fracture and its fracture mechanics extensions(4,5) in which the presence of

mechanically or chemically induced flaws or cracks act as stress concentrators and cause localized failure. Since

the failure stress is inversely proportional to the flaw size, the variability of the strength of optical fibers is a

reflection of the flaw size distribution. For any tested length of fiber, fracture will occur at the position of the

largest flaw. Thus, as the length of the fiber increases, the strength will decrease because of the greater

probability of encountering increasingly larger flaws. Extrinsic flaws can result from contaminants in the preform

or fiber, airborne particles in the processing environment, particle contamination in the polymer coating, or

surface abrasion during processing. There has been much effort to improve the fiber materials and processing

techniques to minimize the number of large flaws in today's high performance fibers. (6,7)

Once the fiber is made, the larger remaining flaws can be eliminated by stressing (proof testing) the entire

length to a level determined by a fatigue design analysis for the particular application. This provides a guaranteed minimum initial strength or a corresponding minimum lifetime for a particular length of fiber. However, when proof testing at high levels, one must be concerned about damage or fatigue which may occur

during the testing. In order to insure a guaranteed minimum strength it is necessary to implement environmental

controls during prooftesting along with rapid unloading of the proof test load.(8,9) Fig. 1 shows a proof tester design which has been used successfully up to levels of 500 ksi (3.5 GPa).(10,11) During proof testing, the fiber

passes between the belts and wheels with 2 turns around the lower floating pulley on which the load is hung.

Proof testing is normally conducted at relative humidities of 40 - 50% with the stress being applied for 1 sec and

unloaded in < 0.1 sec.

In addition to the initial strength degradation due to flaws in the silica fibers, a time dependent phenomenon

known as delayed failure or fatigue can also occur. In this case, a stress lower than that which causes immediate

failure allows the flaws to grow slowly in an active environment until they reach the critical size for fast fracture.

It is generally agreed that a stress corrosion mechanism is responsible in which the stress enhancement of the

reaction SiOSi + H20 —> 2SiOH causes the cracks or flaws to lengthen. This subcritical crack growth is usually

described by the power law model of Charles(12) and the application of fracture mechanics principles(13) which

leads to the relationship between the time to failure, t, and the applied stress, a,

log t = -n log a + log (Ba"~2

)

where (jj is the fracture strength in an inert environment and B and n are parameters related to the material and

environment.

69
The logarithm of the applied stress is usually plotted as a function of the logarithm of the time to failure

typically resulting in a linear plot with a reciprocal slope, n, the stress corrosion susceptibility parameter, and an

intercept at logc = 0 of log (Ba"-2

). This plot is normally extrapolated to make long-term predictions of static fatigue/14' 15) The value of n is typically 20 for silica fibers in non-aggressive environments but can vary widely

due in part to the large scatter in failure times resulting from the statistical distribution of maximum flaw sizes in

repeated tests. Further difficulties arise in predicting fatigue behavior because of the complex interactions of

environmental conditions such as temperature, humidity, pH, and polymer coating chemistry.

There are two types of tests for time dependent failure of silica fibers which are differentiated by the manner

in which the stress is applied. The static fatigue test is conducted in an active environment by applying a

constant stress below that required for fast fracture, measuring the time to failure, and plotting log a vs log t to obtain the stress corrosion susceptibility factor, n. The dynamic fatigue test measures the breaking strength of

the fiber at a constant strain rate in an active environment. By plotting log failure stress vs log strain rate or log

time to failure the n value is also obtained.

There are three techniques, shown on Figs. 2 & 3, commonly used for measuring static and dynamic fatigue:

tension, mandrel bending, and 2-point bending. In the tensile technique,(16) a length of fiber is gripped at each

end and either placed in tension using weights or pulled to failure in a tensile tester. The coated or uncoated

center section of fiber can be threaded through an environmental chamber. However, the coating must remain

intact where the fiber passes through the rubber seal and at the grips to prevent mechanical damage.

In the mandrel bending technique(17) a coated fiber is wrapped around a rod of known radius which defines a

tensile strain on the outside surface. The entire mandrel may be placed in the environment to be studied. The

major advantage of this technique is the ease of preparing many test specimens using rods of different radii. However, there is a serious problem with gripping the fiber ends. Stress concentrations can be introduced at these points causing premature failure. A modification of this technique(18) has been described, however, which

apparendy avoids this problem.

The third technique, two-point bending/1 *"*2^ can be used for either dynamic or static fatigue measurements,

as shown in Fig. 3. The dynamic fatigue test involves constraining a bent loop of fiber between two polished

face plates which are then brought together in a controlled manner until the fiber breaks. The static fatigue test

is conducted by bending a number of fibers and inserting them into a precision bore glass tube or other suitable

fixture as shown in Fig. 3. The stress, in this case, is determined by the inside tube diameter and failure of the

fibers is monitored acoustically. The advantages of these techniques include the elimination of gripping

problems, the ability to easily test coated and uncoated fibers in various environments, and in the case of static

fatigue, the convenience of being able to test many specimens at the same time.

Fig. 4 shows a summary of some static fatigue data obtained from tensile tests and gives an indication of its

sensitivity to temperature, pH, humidity, and polymer coating. It has been found that, under certain

environmental conditions, log t is not linearly dependent on log a with an n value of -20, but shows a transition

in slope resulting in more rapid degradation. This transition was first observed by Wang and Zupko(23) at 32°C

and 90% relative humidity for a fiber coated with an epoxy acrylate. Under these conditions, the value of n

abruptly decreased to -7 after approximately 20 days. Krause's experiments(16,24) showed that the time for the

transition changed to 1 day and 1 hour for coated and uncoated fibers, respectively, tested in water at 90"C.

However, neither the presence of a coating or the increase in temperature changed the post-transition n-value of

~7. Chandan and Kalish(25) later examined the static fatigue of a polyurethane acrylate coated fiber in 404 C - 90°C water and also found enhanced fatigue at low stress levels, with n-values varying from 1-8 depending on

temperature. Although Chandan and Kalish extrapolated the temperature dependence of n and concluded that

there was no enhanced fatigue at temperatures of 3"C, it has been suggested (15,26) that their temperature

dependence was a result of coating effects, and that enhanced fatigue should not be neglected in lifetime

estimates. The coating influence on static fatigue is underscored in Fig. 4 by the absence of a static fatigue

transition for fibers with a two-layer silicone(27) and a fluorinated acrylate(28) coating tested in 80"C and 100°C

water, respectively.

Fig. 4 also describes the dramatic pH dependence of static fatigue failure times for uncoated fiber at 90"C(15)

. The data before the transition shows a dependence of n on pH with n varying from 31 at pH 1, to 13 at pH 14. The post-transition n values, however, are independent of pH, with n =6. Mathewson and Kurkjian(26) show a

generally similar dependence of static fatigue on pH from two-point bending experiments. They also show a

70
crossover in fatigue behavior between coated and bare fibers. At low applied stress coated fiber has the shorter

lifetime while the converse is true at higher stresses. Thus the effect of environment and polymer coating

chemistry on static fatigue is very complex and must be carefully considered when making fatigue measurements

and lifetime predictions.

Because of the permeability of the polymer coatings to moisture and their complex surface interactions which

affect static fatigue, thin metallic(29) or ceramic-like(30-32) hermetic coatings have been deposited on the fiber

surface to improve long-term reliability. Measurements using both tension and bending have shown dramatically

improved static fatigue resistance for some of these hermetically coated fibers. For example, a design analysis

for a 25 year service fife of a carbon coated fiber shows that the service stress can be -80% of the prooftest

stress, as compared to -30% normally estimated for conventional fiber.(33)

HYDROGEN

Optical fibers are known to experience detrimental loss increases if they are exposed to sources of hydrogen.

H2 can readily diffuse into a fiber's core, causing loss increases by one of two basic mechanisms: (1) H2 becomes IR active when dissolved in a silica glass matrix, giving rise to absorption peaks between 1.08 and

1.26|j.m, and a rising loss edge beyond about 1.4|am (34,35)

, (Fig. 5) and (2) H2 reacts with structural defects in the

glass, causing, for instance, the formation of OH at nonbridging oxygen sites in the fiber core.

(36,37) The physical

dissolution of H2 molecules in Si02

is a reversible process. Therefore these loss increases will gradually

disappear if the H2 atmosphere is removed. Losses due to this dissolved H2 are substantially the same for all

fibers, regardless of the glass composition. On the other hand, the chemical reaction of H2 with defects generally

results in permanent loss increases, the magnitudes of which depend sensitively on the glass composition/38,39*

The issue of hydrogen reliability is of concern because there is always some finite level of H2 present in a

cable, due both to H2 outgassing of polymers and to possible galvanic corrosion of cable materials. One strategy

is to avoid hydrogen problems by means of suitable cable design, or by the use of non-reactive or hermetic

fibers. (31,32' 39) However, it is frequently necessary to characterize a given fiber type so as to allow the accurate

prediction of long term loss increases, and to permit the specification of a "tolerable" hydrogen level within a

cable. The remainder of this section will discuss the identification of hydrogen induced losses and the prediction

of fiber loss increases in hydrogen-containing environments.

Determining that hydrogen induced losses have indeed occurred is best done by use of spectral loss

measurements, preferably taken before and after the occurrence of the loss increases. Molecular hydrogen has a

distinct spectral signature, as shown in Fig. 5. The magnitudes of these losses are proportional to the

concentration of H2 in the fiber core. The rate of loss increase is determined by the diffusion rate of H2 in

SiO2

(40) and is readily predicted using classical diffusion theory. (41) The hydrogen level in the core of a single

mode fiber will equilibrate with an applied atmosphere in about 2 weeks at 21°C. Thereafter the losses due to

this effect will be constant, as long as the external hydrogen level does not change. A fiber exposed to 1 atm of

H2 at 21°C will show loss increases of about 0.21 and 0.55 dB/km at 1.31 and 1.55|im respectively, and a 9.5

dB/km loss increase at the 1.24)j.m H2 overtone peak.

The formation of OH at defect sites causes readily measured increases in the OH overtone in the 1.4(im

region. If sufficient spectral resolution is used, it will frequently be noted that the peak has two components,

centered on 1.39 and on 1.42|im, indicative of SiOH and GeOH, respectively.(42) In the case of fibers co-doped

with Ge and P the growth of the GeOH component can be quite pronounced, and is accompanied by a rising

long wavelength loss edge.(36,39,43) Qualitatively similar behavior has been noted in single mode fibers

contaminated with unusually high levels of alkali impurities.(44,45) In Ge-doped single mode fibers the OH

overtone growth occurs primarily at 1.39|im, but can have a detectable 1.42|im shoulder. (46)

Reaction of H2 with other types of defects in Ge-doped fibers causes a short wavelength loss edge that can

significantly influence losses at 1.31 and 1.55 |im, but which is most readily noted in the wavelength region

below l^im. (Fig. 5)

(47,48) The absorption cross section of the defect responsible for this edge is temperature

dependent, and thus the magnitude of the excess loss depends on the measurement temperature.(49) This loss edge

is often the dominant hydrogen loss mechanism for Ge-doped single mode fibers at typical operating conditions.

71
Predicting losses due to molecular H2

is quite straightforward if the hydrogen pressure is known, as previously discussed. However, in many cases the pressure of H2 around a fiber is rather low, typically below

10~3 atm. Molecular H2 peaks will not be observable under these conditions, and reactive types of loss increases

will occur very slowly. The concern in these cases is that small but "significant" loss increases might occur over prolonged periods, due to reactions at defects. In order to predict the loss increases it is necessary to accelerate

the reactions by raising the temperature and/or the hydrogen pressure, and then to extrapolate back to anticipated

temperatures and pressures. To make valid predictions it is necessary to separately characterize the temperature

and pressure dependences of each type of hydrogen induced loss. In practice, one type of loss will often be

dominant in a given fiber type, (e.g. the short wavelength edge in Ge-doped single mode fibers), and it is this mechanism that must be carefully characterized. The results obtained for one type of fiber (e.g. a Ge-doped

single mode fiber) will usually not be applicable to other fiber types (e.g. Ge-P doped multi-mode fibers).

The use of in-situ spectral loss measurements is invaluable in characterizing the reactions that occur in these

accelerated tests.

(47) Monitoring the loss changes of a fiber while it is exposed to H2 at an elevated temperature

will show an initial rapid rate of increase due to molecular hydrogen, followed by more gradual increases due to reaction of H2 at defects. (Fig. 6) The reactive component of the loss increase itself has an initially rapid rate of

increase, followed by a slow approach to a saturation level as reactive defects are depleted. (47) As long as only a small fraction of the defects have been depleted by reaction it is fair to approximate the rate of loss increase to be a constant, resulting in loss changes that are proportional to the exposure time. It is therefore, the initial rate of loss increase that is generally characterized, usually at several temperatures and hydrogen pressures for a given

fiber type. Ideally, an Arrhenius relation should be followed for experiments run at different temperatures and a given hydrogen pressure. Fig. 7 shows such behavior for Ge-doped single mode fibers.(47,49) The straight line fit indicates that a single mechanism is dominant over the temperature range studied. Extrapolation to lower

temperatures necessarily assumes that a lower slope portion will not occur on the Arrhenius plot at lower

temperatures. For this reason it is best to perform experiments at temperatures as close as possible to actual

operating temperatures, although this will entail experiments of longer duration.

Similarly it is necessary to determine the dependence of the initial reaction rate on the hydrogen pressure.

Results to date for several fiber types have shown that the rate of loss increase is usually proportional to the

square root of the hydrogen pressure at relatively high pressures (Ph2 > 0.1 atm).(39,50) However, experiments on Ge-doped single mode fibers have shown that the dependence on the hydrogen pressure changes to a linear

dependence at low hydrogen pressures (PHz < 0.04 atm).(47,49) Using this data for the temperature and pressure

dependences it is possible to determine tolerable hydrogen levels for given values of allowable loss increases, as shown in Fig. 8.

(47) At low temperatures, reactions are slow and the losses due to molecular H2 are the primary

effect. At higher temperatures the hydrogen reactions are faster and the loss increases are primarily due to the

growth of a short wavelength loss edge. The loss increases due to OH, while noticeable near 1.4|j.m, have little

effect at 1.31 and 1.55|j.m.

Comparison of two different types of fiber, A and B, using a "standard" accelerated test at some selected

temperature and hydrogen level requires some caution. A meaningful comparison experiment requires knowledge

of both the temperature and the pressure dependences for each fiber. For instance, assume that rates of loss

increase in fiber A have a stronger temperature dependence than those for fiber B. At high temperatures, loss

increases will be larger for fiber A than for B, while at low temperatures the converse will be true. (That is, the

lines for A and B will cross on an Arrhenius plot.) A relative ranking of the two fibers would therefore depend on the temperature chosen for the experiment. A similar argument can be made regarding the importance of

choosing a meaningful hydrogen pressure for a comparison experiment.

Recent advances have made it possible to apply hermetic carbon coatings to optical fibers.

(31,32) These

coatings dramatically reduce the levels of H2 seen by the lightguiding portion of a fiber, and allow the fibers to be used in applications where the external hydrogen pressure is quite high. Because so little H2 penetrates these

coatings at normal conditions it has been necessary to use both elevated temperatures and high Ph2 's to get

measurable loss changes in accelerated tests. Pressures of 2000 psi and temperatures from 120 to 200°C lead to

detectable H2 permeation, making it possible to estimate the permeation time constants for real operating

conditions. (51)

72
RADIATION

The loss added to an optical fiber by interactions with ionizing radiation is of interest in both low dose

terrestrial and undersea environments as well as the high dose surroundings of a nuclear blast, nuclear reactor or

spacecraft/

52-56

* Prediction of the fiber response in these surroundings is complicated by the substantial variations

in total dose, dose rate, temperature and radiation type, as evident from Table I. Correspondingly, the

measurement and prediction of fiber responses must take these variables into consideration. Nuclear blast studies

frequently consider transient components which require specific methods not discussed here.(57,58) Fiber radiation

sensitivities are further complicated by dependence on wavelength, mode excitation, launched optical power and

ambient light. Attempts to establish a standard testing procedure have recently been made by the NATO Panel

IV, Research Study group 12 and by the Electronics Industries Association.*

59-61 *

Steady state radiation induced loss measurements, typically used to qualify fiber, are divided into two ranges,

low dose (<3000 rads at a dose rate of 5 to 18000 rads/hr) and high dose (3000 to 1,000,000 rads at a dose rate of 5 to 200 rads/sec) with low dose measurements made using a before-and-after cutback method. In high dose

measurements the loss is monitored throughout the exposure and for several minutes following the

exposure.

(59-61)

Common to all radiation measurements is the need for accurate dosimetry. The radiation dose is typically

expressed in rad (Si) where one rad equals an absorbed energy of one erg per gram of silicon. Standard methods

to measure the absorbed dose employ thermal luminescent detectors (TLD's), radiochromic dyes (used by NBS),

and Fricke dosimetry/62

* The TLD method is perhaps the most convienient in that the detectors may be

purchased as chips, a few millimeters square, which are reusable after annealing and may easily be placed on the

fiber reel. The TLD material is chosen depending on the dose range of interest, with lithium fluoride suggested

for the range of 1 mrad to 300,000 rads and lithium borate manganese for the range of 50 mrad to 1,000,000

rads, to name a few.(63) Reading the TLD is readily accomplished using commercial equipment which integrates

the light output of the material as it is subjected to a well controlled temperature cycle. This output reading

serves only as a relative indication of the absorbed dose and requires comparison with the integrated light output

from a known dose and verification of linear response with dose for calibration. The TLD materials respond

independent of dose rate (up to 1012 rad/sec.) and energy (from less than 10 kev to several Mev) and fade less

than 5% per year at room temperature. A reproducibility of approximately 20% in the measured dose can be

expected with this technique. Mixed gamma/neutron doses may be measured with combinations of TLD's. The

Fricke dosimetry method is based on the change in optical absorbance at 304 nm by ferric ions created during the

radiation of a ferrous ammonium sulfate aqueous solution. This has been a widely accepted chemical dosimeter,

as have the dyes, with claimed accuracy of ±5%. However, the Fricke dosimeter is limited to a relatively small

dose range of 4 to 200 krad by the oxygen content of the solution.

Control of dose uniformity over the area of the fiber reel is an additional concern given that radiation sources

are typically contained in single or multiple point source configurations. Dose field uniformity may be achieved

by placing the fiber reel sufficiently far from the source. However, this also decreases the dose rate. One

manufacturer of a Co60 source achieves a flat dose field by providing a variable thickness lead shield to compensate for differences in dose rate over the sample area/65

* Dose non-uniformity over a reel of fiber may

also arise from shielding by overwound fiber, a potential problem in the low dose, several kilometer length

sample configurations.

Equipment and procedures for measuring the radiation induced losses in fibers are typically adaptations of

standard loss test sets as shown in Figure 9, with measurement procedures dependent upon the total dose and

dose rates of the exposure.

Evaluation of the low dose (<3,000 rad)/low dose rate (5 to 50 rad/hr) sensitivities is performed using the

standard cutback technique, before and after the irradiation. Care must be taken to ensure that the ambient light

is not contributing to photobleaching both during exposure and during measurement. Fiber lengths of a few

kilometers are required to observe a few tenths of a dB added loss in state of the art radiation resistant fiber. Low dose (< 3,000 rad)/high dose rate (50 to 20,000 rad/hr.) and higher dose measurements are usually

performed in-situ, monitoring the added loss both during and after exposure. Measurements of this type for three

generic fiber compositions are shown in Fig. 10, demonstrating the diversity in responses/64

* As a further

complication, these responses have a temperature dependence as shown in Fig. ll.(65) In these tests, less than 100

73
meters of fiber are used and control of an equilbrium mode distribution becomes more important to insure

repeatability. A stable source and launch are also required in that the exposure may last an hour, with recovery

typically monitored for nearly 3 hours.

In general, the optical power present during the measurement has been shown to affect the induced losses,

with the higher power causing lower induced losses. This photobleaching effect is especially noticable in

undoped silica core fibers. However, the average optical power during in-situ measurements of any fiber type is suggested to be l|iW or less to minimize the effect.

(59,60,61) The photobleaching effect may be further minimized

by sampling the transmission only a few times during the measurement/61

*

In summary, the measurement of the radiation response of an optical fiber must be designed for the specific

environment of interest, and is established primarily by total dose, dose rate, temperature and radiation type. The

fiber response to this environment is, in turn, a complex sum of effects mainly determined by its composition, the

wavelength, and the optical power of the probe signal.

CONCLUSIONS

The three relibility issues discussed here are associated with degradation mechanisms affected by complex

interactions of materials and environment. Inherent difficulties of extrapolation for predicting long-term

reliability necessitate a thorough understanding of the degradation processes in order to allow the design of

meaningful test methods.

REFERENCES

1. S. R. Nagel, SPIE Proceedings, 717, 8-20, Sept. 1986.

2. B. A. Proctor, I. Whitney, and J. W. Johnson, Proc. R. Soc. London, Ser. A. 297A, 534-557, 1967.

3. A. A. Griffith, Phil. Trans. Roy. Soc, London, A 221, 163, 1920.

4. G. R. Irwin, Encycl. Phys. 6, 55, 1958.

5. S. M. Wiederhorn, J. Am. Ceram. Soc. 52, 99, 1969.

6. F. V. DiMarcello, A. C. Hart, J. C. Williams, and C. R. Kurkjian, Physics of Fiber Optics: Advances in

Research and Development, ed. B. Bendow and S. S. Mitra, Plenum, N. Y. 125-135, 1979.

7. A. L. Harmer, Top Meet. Opt. Fiber Commun., Tech. Dig WF5, 70-73, 1979. 82-85, 1979.

8. A. G. Evans and E. R. FuUer, Mater. Sci. Eng. 19, 69-77, 1975.

9. B. K. Tariyal and D. Kalish, In "Fracture Mechanics of Ceramics", Vol 3, R. C. Brandt, D. P. H. Hasselman,

and F. F. Lange, eds. Plenum, N. Y. 161-195, 1978.

10. Private Communication, J. G. Wright and D. L. Myers, AT&T Technologies, Norcross, GA.

11. F. V. DiMarcello, D. L. Brownlow, R. G. Huff, A. C. Hart, Proceedings OFC '85, PD06, San Diego, 1985.

12. R. J. Charles, J. Appl. Phys. 29, 1554-60, 1958.

13. A. G. Evans, Int. J. Fract. 10, 251, 1974.

14. J. T. Krause, D. A. Meade, S. Shapiro, Proc. Suboptic Int. Conf. on Opt. Fib. Subm. Telecomm. Sys., Paris,

117-22, 1986.

15. J. T. Krause and C. J. Shute, Adv. Ceram. Mat. 3, 118-121, 1988.

16. J. T. Krause, J. Non-Cryst. solids, 38-39, 497-502, 1980.

17. M. Nishimura, K. Yoshida, M. Oda, Y. Ishida, Proceedings of 6th ECOC, York, 133-136, 1980.

18. P. C. P. Bouten and H. H. M. Wagemans, Electron Lett. 20, 280-81, 1984.

19. P. W. France, M. J. Paradine, M. H. Reeve, & G. R. Newns, J. Mater. Sci. 15, 825-30, 1980.

20. S. F. Cowap and S. D. Brown, Am. Ceram. Soc. Bull. 63, 495, 1984.

21. M. J. Mathewson, C. R. Kurkjian, S. T. Gulati, J. Am. Ceram. Soc, 69, 815-21, 1986.

22. M. J. Mathewson, C. R. Kurkjian, J. Am. Ceram. Soc, 70, 662-68, 1987.

23. T. T. Wang and H. M. Zupko, J. Mat. Sci. 13, 2241-48, 1978.

24. J. T. Krause, Proceedings of 5th ECOC, postdeadline paper, 19.11-19.4, Amsterdam, 1979.

25. H. C. Chandon and D. Kalish, J. Am. Ceram. Soc. 63, 171-73, 1982.

26. M. J. Mathewson and C. R. Kurkjian, J. Am. Ceram. Soc. 71, 177-83, 1988.

27. D. Roberts, E. Cuellar, L. Middleman, and J. Kucker, SPIE Proceedings, Paper 721-14, 1986.

74
28. B. J. Skutnik, B. D. Munsey, and C. T. Brucker, Mat. Res. Soc. Sym. Proc. 88, ed. S. R. Nagel, J. W.

Fleming, G.H. Sigel, and D. A. Thompson, MRS, Pittsburgh, 1987.

29. D. A. Pinnow, G. D. Robertson, G. R. Bair, and J. A. Wysocki, Top. Meet. Opt. Fiber Commun. Tech. Dig.

TUCS, 16-17, 1979.

30. E. G. Hanson, et al, U.S. Patent 4, 512,629, April 23, 1985.

31. R. G. Huff, F. V. DiMarcello, and A. C. Hart, OFC '88 TUG2, New Orleans, 1988.

32. K. E. Lu, G. S. Glaesemann, R. V. Vandewoestine, and G. Kar, Journ. Lightwave Tech 6, 240-43, 1988.

33. J. T. Krause, C. R. Kurkjian, F. V. DiMarcello, and R. G. Huff, 6th EFOC/LAN Conf., Paper 2.8

Amsterdam, June 29 - July 1, 1988.

34. J. Stone, A. R. Chraplyvy, C. A. Burrus, Optics Letts., 7, 297, 1982.

35. K. Mochizuki, Y. Namihira, Electron. Lett., 19, 743, 1983.

36. N. Uesugi, T. Kuwabara, Y. Koyamada, Y. Ishida, N. Uchida, Appl. Phys. Lett., 43, 327, 1983.

37. P. J. Lemaire, M. D. deCoteau, Optical Fiber Materials and Properties, Materials Res. Soc. Symp. Proc, 88,

225, MRS, 1987.

38. K. Noguchi, N. Shibata, N. Uesugi, Y. Negishi, J. Lightwave Tech., LT-3, 236, 1985.

39. J. D. Rush, K. J. Beales, D. M. Cooper, W. J. Duncan, N. H. Rabone, British Telecom Technol. J., 2, 84,

1984.

40. R. W. Lee, R. C. Frank, D. E. Swets, J. Chem. Phys., 36, 1062, 1962.

41. J. Crank, The Mathematics of Diffusion, Oxford Univ. Press, London, 1975.

42. J. Stone, G. E. Walrafen, J. Chem. Phys., 76, 1712, 1982.

43. S. Tanaka, M. Kyoto, M. Watanabe, H. Yokota, Electron. Lett., 20, 283, 1984.

44. A. lino, J. Tamura, K. Orimo, K. Okubo, Conf. on Optical Fiber Comm, Atlanta, 1986. Paper TU13.

45. H. Itoh, H. Hanafusa, Y. Hibino, Trans. IECE Japan, E69, 782, 1986.

46. P. J. Lemaire, A. Tomita, European Conf. Optical Comm., Stuttgart, 1984. Paper 15A1.

47. A. Tomita, P. J. Lemaire, Electron. Lett., 21, 71, 1985.

48. N. J. Pitt, A. Marshall, Electron. Lett., 20, 513, 1984.

49. A. Tomita, P. J. Lemaire, European Conf. Optical Comm., Stuttgart, 1984. Postdeadline paper PD-1.

50. K. Noguchi, N. Shibata, N. Uesugi, K. Ishihara, Y. Negishi, Electronics and Commun. in Japan, 69, 1, 1986.

(Translation from Denshi Tsushin Ronbunshi, 68B, 795, 1985.)

51. P. J. Lemaire, K. S. Kranz, K. L. Walker, R. G. Huff and F. V. DiMarcello, Proc. of 14th Europ. Conf. on

Optical Comm., Brighton, UK, 1988.

52. J. B. Haber, E. Mies, J. R. Simpson, and S. Wong, J. Lightwave Tech. 6, 150-154, 1988.

53. H. J. Schulte, Proceedings of OFC, paper TUQ2, San Diego, 1985

54. R. E. Warren and A. Beilis, Report prepared for the Defense Nuclear Agency under contract No. DAEA-18-

75-A-0059-8Z01AC by AT&T Technologies, 1985.

55. J. K. Partin, Proceedings of SPIE Fiber Optics in Adverse Environments, 506, 42-49, 1984.

56. E. J. Friebele, Proceedings of OFC, paper WA7, Reno, 1987.

57. L. D. Looney, G. Turquet de Beauregard, P. B. Lyons and R. E. Kelly, Proceedings of SPIE Fiber Optics in

Adverse Environments, 296, 17-25, 1981.

58. E. J. Friebele, J. Lightwave Tech., to be published.

59. Electronics Industries Association (EIA), Wash. D.C., "Proposed Procedure for Measuring Gamma

Irradiation Effects in Optical Fibers", FOTP-49, 1987.

60. R. A. Greenwell, Fiber and Integ. Opdcs, 6, 179-80, 1987.

61. E. J. Friebele, E. W. Taylor, G. Turguet de Beauregard, J. A. Wall, and C. E. Barnes, J. Lightwave Tech. 6,

165-71, 1988.

62. J. H. O'Donnell and D. F. Sangster, "Principals of Radiation Chemistry", Amer. Elsevier Publishing Co.

Inc., New York, 39- 41, 1970.

63. Harshaw/Filtrol, TLD systems & materials performance specifications

64. J. R. Simpson, Proceedings of OFC'88, Paper TU-G3, New Orleans, 1988.

65. J. R. Simpton, T. Mensah, S. Wong, A. J. Ritger, E. Mies, A. Klein and P. Glodis, Proceedings of ECOC,

Helsinki, 1987.

75
Fig 1. Constant load prooftester. Fig 2. Fatigue test methods.

a. DYNAMIC b. STATIC

GUIDE PLATE

GLASS TUBE

FIBERS

Fig 3. Two-point bend techniques.

•jo 500 - w 400

CO

LU

F 300

.KRAUSE

pH 1 90°C

i ROBERTS

(2- LAYER SILICONE)

80°C

10 2 103 104 10 5 106 107

TIME TO FAILURE (sees)

Fig 4. Static fatigue of silica fiber.

76
X(/u,m)

Fig 5. Loss spectra for untreated and for Fig 6. Loss increases versus time for fiber

hydrogen treated fibers, showing basic hydrogen exposed to hydrogen, showing components due

induced losses. to molecular H2 and due to reactions at defects.

T(°C)

0 250 200 150 100 75 50 25 0

1.5 2.0 2.5 3.0 3.5 4.0

1000/T(K" 1

)

Fig 7. Temperature dependence for initial rate of loss increase in Ge-doped single mode fibers.

Short wavelength loss edge component measured at l.O^im. Hydrogen pressure is 1.0 atm.

0 20 40 60 80 100

T(°C)

Fig 8. Tolerable hydrogen pressures for Ge-doped

single mode fibers. Curves are for different values

of allowable loss increases at 20 years.

TUNGSTEN

HALOOEN

LKJMT SOURCE

II

LOCK-M

AMPUPKR

CHART

RECORDER

OPTICAL MECHANICAL PITERS / OPTICAL CHOPPER MON0-

CHROMETXR

•PUTTER

INPUT LAUNCH

SIMULATOR

CLADDING

MODE

•TRVPCR

OPTICAL DETECTOR

TEMPERATURE CONTROLLED

CHAMBER

REEL OP

TEST

PISER

zzzzzz ' SHUTTER

Fig 9. Instrumentation for testing the effects

of adverse nuclear environments.

•HIELOED CHAMBER

OAMMA

RADIATION SOURCE

77
Fig 10. Radiation induced loss responses for three

common optical fiber core compositions: a, silica; b, germanium doped silica; c, phosphorus doped silica.

EXPOSURE RECOVERY

SMI - 0.6 WT % P CLADDING 6M2 • LOW EFFECTIVE P CLADDING

-55'C RESPONSES

SMI

I I I I I I

• x 10 RAD CO

200 RAD/SEC

f 2S0 FT

I

w2

S

oa

<

300 600

TIME (SEC.) -40 -20 0 10 20 30 TEMPERATURE *C

Fig 11. Temperature dependence of the

radiation induced loss for two single mode

fibers.

RADIATION ENVIRONMENTS

Environment Radiation Type Dose Rate Temperature

Terrestrial Gamma from

Uajt Kttt TII232

100-500

mrad/yr

-50-150*C

Undersea Gamma from

Ujjj, K40. TI1232

100

mrad/yr

3*C

Nuclear Blast Mixed gamma, neutrons and charged particles

worst case

free field at 2 psL

overpressure 35,000 rads at 16 rads/sec

-50-200*C

Nuclear Reactor Mixed gamma, thermal and fast neutrons gamma

10-10* rad/hr/MW

neutron 1(T5 -1012 n/cm2 -sec/MW

-150-120'C

Spacecraft Mixed gamma and

charged particles

1-46 rad/day -150-120'C

Table 1 Radiation Environments

78
FLUORIDE OPTICAL FIBER MEASUREMENTS: TECHNIQUES AND STATUS

Lynda E. Busse

Naval Research Laboratory

Optical Sciences Division

Washington, D.C. 20375

From the time when the first fluoride glasses were made in

1974 and it was determined soon thereafter that these glasses

had the potential for losses lower than silica in the infrared,

there has been a strong research effort at laboratories

throughout the world to produce optical fibers from these

materials [1-3] . Calculations and extrapolations based on

experimental data of the intrinsic Rayleigh scattering level

and the contribution from the multiphonon edge gave a determination of a minimum loss "window" to be near 2.5 microns with

a loss of 0.01 dB/km [2]. The interest is in obtaining long

lengths of low loss fluoride fibers for optical communications

applications

The nature of fluoride glasses causes their fiberization

to be difficult due to the intricate and interwoven problems

which must be overcome. The glass compositions typically

contain five or six heavy metal fluoride components, for

example ZrF4-BaF2-LaF3-AlF3-NaF . Each of these components must

be purified of transition metal and rare earth elements which

absorb in the infrared. In addition dry preparation conditions

must be maintained to reduce the OH absorption loss at 2.9

microns. While silica is inherently stable to crystallization

during melting, fluoride glasses are very unstable upon cooling

or reheating. Thus process control must be stringent to

prevent the presence of defects such as crystals or bubbles in

the fibers [1-3]

Control of waveguide parameters is also essential. The

fibers have a step index profile which must be accurate, and

core diameter control and core/cladding concentricity must be

maintained. To date fluoride fibers have been mostly multimode, with core diameters of 50-100 microns and numerical

apertures from 0.1 to 0.2. Some recent fibers which are single

mode will be discussed.

In this paper fluoride optical fiber measurement techniques will be discussed, including those which are typical of

conventional silica measurements (e.g. cutback and scattering)

as well as some other types which have been developed recently

to give insight into the sources of defects and absorption

loss. The experimental techniques and current loss levels will

be discussed.

TOTAL LOSS To measure the total loss in a given length of

fluoride fiber, the cutback method, similar to that used for

silica fibers, has been used. There are some differences in

the experimental setup, however. At NRL , a high intensity

infrared zirconia source is used; the light is directed through

a set of order-sorting filters and into a monochromator which

has a computer-driven stepper motor. Two gratings are used for

79
scans from 1-4 microns, with step size of 0.01 micron. The

light exiting the monochromator is chopped and then collimated

and focussed by two spherical mirrors, mounted slightly offaxis. The light is focussed to match the fiber numerical

aperture, and the spot size at the focus is either equal to or

greater than the core diameter. The fiber output end is mode

stripped by immersing it in index matching fluid. The light

exiting the fiber is collected by an indium antimonide infrared

detector which is connected to a lock-in analyzer interfaced to

the computer. Care is taken to do the measurement at zero

tension

Figure 1 is a summary plot [2] of the best losses

achieved to date at British Telecom Research Laboratories,

Nippon Telephone and Telegraph (Japan) , and NRL. NRL and NTT

have achieved 1 dB/km [4] and 0.7 dB/km [5], respectively, over

30 m. lengths. BTRL has achieved 3.9 dB/km over 100 m. as

shown and 9.5 dB/km over 400 m. [6]. Due to the inhomogeneous

number of scattering defects in the fibers, longer lengths

often show higher loss. The consistency of total loss results

among the different laboratories has not been fully tested, but

will be essential as losses get even lower. Several kilometers

of fiber will be needed to accurately determine losses as they

approach the theoretical limits.

Some single mode fluoride fiber results have been reported

[7,8]. The lowest loss achieved was 3 dB/km at 2.5 microns

[8] . The core diameter was 11 microns, An=0.47%, and the

cutoff wavelength was 2.1 microns.

SCATTERING LOSS Since the total loss is composed of

scattering and absorption, it is necessary to determine each to

evaluate a fiber's quality. There are many sources of extrinsic scattering loss in fluoride fibers, including crystals,

bubbles, sub-micron scattering centers, and interface defects.

Tighter process control has substantially lowered the number of

crystals and sub-micron particles, but defects still exist.

Thus an accurate determination of the scattering loss in

fluoride fibers is essential for process improvement.

The measurement is done with a scattering sphere detector,

similar to work done on silica fibers over ten years ago. The

loss is calculated by: <X(dB/km) = 4.34 x 10 5" [P Sc /(P© x

L) ] , where P sc = Scattering intensity; P 0 = Intensity of light

exiting the fiber; and L = Diameter of the scattering sphere.

Several measurements have been made in the visible with a

krypton laser [9], for example, using a scattering sphere which

has an inner white reflective surface. A plot was made to

determine whether the scattering follows a Rayleigh scattering

law (i.e. varies as ) or Mie scattering (i.e. varies as

X*) . The scattering value at 2.5 microns is then inferred from

an extrapolation.

At NRL scattering measurements can be done directly at 2.5

microns with a color center laser. The scattering sphere is

gold coated to be reflective in the infrared, and an indium

antimonide detector is used. Although only 5 cm. lengths are

measured at a given time, many such measurements can be made to

80
give a general level of scattering in the fiber.

Although present fibers are of good quality such that the

losses obey a Rayleigh scattering behavior, there is a nontrivial contribution to the scattering due to a wavelength

independent offset. Early fluoride fibers had wavelength

independent offset values of several dB/km [10] but more recent

fibers show an offset of less than 1 dB/km [9,11]. Figure 2

shows such an offset for fluoride fibers measured in the past

at NRL. The source of this offset has not been determined, but

is suspected to be due to particles of size greater than the

probe wavelength, or core diameter fluctuations.

Recently NRL has reported scattering measurements on a

fluoride fiber which approach the theoretical limit of Rayleigh

scattering [11]. The results are listed in Table I; this fiber

had a core diameter of 30 microns and numerical aperture of

0.13. Note that the measured fiber scattering value at 2.55

microns was 0.025 dB/km, which is almost equivalent to what is

obtained from a Rayleigh scattering extrapolation of the value

(7.5 dB/km) measured at 0.63 microns. Thus there is no

observed wavelength independent loss in this fiber.

There have been other measurements done to probe the

causes of scattering in fluoride fibers. At NTT a threedimensional scattering experiment was devised to give a full

angular scattering plot of a small segment of fiber [12] . A

comparison of a silica fiber and two fluoride fibers of

different composition was made; it was found that the silica

fiber and one of the fluoride compositions gave a Rayleigh

scattering polar plot, while one of the fluoride compositions

showed some backscattering which was attributed to phase

separation

Determining what the scattering defect is poses a difficult problem, as optical microscopy cannot resolve particles

of micron size or smaller. At British Telecom Laboratories,

theoretical calculations were made to determine the effect on

the scattering of various types and sizes of particles suspected to be in their fibers [13] . NTT has developed a direct

technique of analyzing particles in fibers by micro-Raman

scattering [14] . They found that particles in their fibers

gave Raman spectra which matched that of zirconium oxide

crystals. At NRL various sources of scattering have been found

by isolation of the particles and subsequent electron microscopy and diffraction measurements. Fluoride crystals nucleated

by platinum, phosphide impurities, and phase separation are

some of the defects found by such analyses [15]

ABSORPTION LOSS It is also important to measure absorption

loss in fluoride fibers, since there are many infrared absorbing impurities in the raw materials which make the glass; these

include the transition metal ions of iron, cobalt, copper, and

nickel. Purification of the raw materials is a challenging

task, as levels must ultimately reach parts per trillion for

the fiber to be ultra low loss [1] . To measure the absorption

loss, the technique of laser calorimetry as developed by White

and Midwinter [16] for silica glass and fiber has been used

81
for fluoride fibers at various wavelengths using a krypton

laser [17] . The technique involves measuring both the rise

time and temperature increase in a fiber after the laser light

is launched into it. Using known thermal properties of the

glass and fiber parameters, and after calibration with a source

of known heat loss, the absorption coefficient at that wavelength can be determined. As shown in Figure 4, with enough

measured absorption (as well as scattering) values at given

wavelengths , the total loss can be deconvoluted to give the

various impurity levels.

Another critical source of absorption loss is neodynium, a

major contaminant in lanthanum; its absorption peak is at 2.51

microns, which is near the fiber's minimum loss wavelength. A

technique for measuring this absorption using photoluminescence

has been developed at both NRL and GTE Laboratories [18,19].

By using a proper excitation wavelength (e.g. 514 nm.), the

luminescence spectrum may be obtained. Using standards of

known impurity content, the quantity of neo.dynium in a fiber

can be determined [20]

SUMMARY The types of optical measurements made on fluoride

fibers have been discussed. As losses become lower and

approach theoretical limits, there is need for longer lengths

of fiber and comparison of results among laboratories to insure

accuracy of total loss measurements. Scattering measurements

are necessary to improve fiber preparation techniques. For the

first time, NRL has reported scattering values approaching the

theoretical Rayleigh scattering limits. Techniques for

measuring absorption in fluoride fibers have been described,

including laser calorimetry and photoluminescence. Both

scattering and absorption techniques will need to be improved

in order to measure the ultra low transmission losses possible

with these new fluoride fibers.

ACKNOWLEDGEMENT The work of George H. McCabe in obtaining the

NRL scattering data reported herein is gratefully acknowledged.

REFERENCES

1. D.C. Tran et al .

, J. Lightwave Tech., V. LT-2, p. 566

(1984); see also Materials Science Forum, V. 5-6 (1985)

and V. 19-20 (1987)

2. P.W. France et al., Br. Telecom Technol . J., V. 5 (1987).

3. Proc. of the 5th Intl. Symp . on Halide Glasses, Japan,

June 1988, to be publ . in Mat. Sci. Forum.

4. D.C. Tran et al .

, SPIE V. 618, Infrared Opt. Mater, and

Fibers IV, p. 48 (1986)

5. T. Kanamori and S. Sakaguchi, Jap. J. Appl . Phys . , V. 6,

p. L468 (1986)

82
6. P.W. France et al .

, Halide Glasses for Infrared Fiberoptics, ed. by R. M. Almeida, Martinus Nijhoff Publ .

,

p. 253 (1987) .

7. Y. Mimura, Mat. Sci . Forum, V. 19-20 (1987).

8. Y. Ohishi et al., J . Lightwave Tech., V. LT-2, p. 593

(1984), and Y. Ohishi et al .

, Electron. Lett., V. 22,

p. 1034 (1986) .

9. P.W. France et al .

, Mat. Sci. Forum, V. 19-20, p. 381

(1987) .

10. D.C. Tran et al., Electron. Lett., V. 19, p. 165 (1983).

11. I. Aggarwal et al., "Scattering Losses in Fluoride

Optical Fibers", Ref. 3.

12. H. Hattori et al . , Appl . Optics, V. 25, p. 3549 (1986);

H. Hattori et al . , ibid., V. 26, p. 650 (1987).

13. M.W. Moore, "Scattering Losses in Fluoride Glass Fibre",

Ref. 3.

14. T. Kanamori, Mat. Sci. Forum, V. 19-20, p. 353 (1987).

15. G. Lu and J. Bradley, ibid., V. 6, p. 551 (1985); V. 19-

20 (1987); and "Phase Separation and Crystallization in

a Lead-Containing Fluorozirconate Glass", Ref. 3.

16. K.I. White and J.E. Midwinter, Opto-Elec tronics , V. 5, p.

323 (1973)

17. M.W. Moore et al . , Proc . 12ECOC, Barcelona, p. 299 (1986)

18. J. A. Freitas et al., Mat. Sci. Forum, V. 19-20 (1987).

19. W.J. Miniscalco et al .

, ibid.

20. J. A. Freitas et al .

, Ref. 3.

TABLE I

Scattering losses in bulk glass and short lengths of fiber

SCATTERING LOSS (dB/km)

A (microns) LAUNCH N.A. MEASURED EXTRAPOLATED

0.63 (Bulk) — 3.9 0.015

0.63 (Fiber) 0.10 7.5 0.028

2.55 (Fiber) 0.10 0.025

83
0.5 1.0 1.5 2.0 2.5 3.0 3 : 4.0

wavelength, fjm

FIGURE 1 : Comparison of total loss results on fibers made

at BTRL, NTT , and NRL [2]

FIGURE 2

SCATTERING

LOSS

(DB/KX)

WAVELENGTH (MICRONS)

2.55 1.06 0. 63

offset:

typically

0.3 - 0.5

dB/km <

1 2 3 4 5 6

WAVELENGTH "4 [ (MICRONS) ~4]

Past NRL fiber scattering results, showing

the wavelength independent offset [11]

0.45 0.55 0.85 0.95

FIGURE 3

0.65 0.75

wavelength, pm — Deconvolution of loss to obtain impurity levels [17]

84
BEND IPSS MEASUREMENTS ON SINGLE MODE FIBERS

A. Hordvik, M. Eriksrud, S. Lauritzen

Optoplan, P.O.Box 1963 N-7002 Trondheim, Norway

Introduction

The bend loss sensitivity of a fiber is of considerable practical importance.

There often is a need to wind short lengths of fibers in a small coil f.inst.

in splice boxes, and in cables where fiberbending is an important loss

mechanism [1,2]

This paper presents the results of bend loss measurements on a number of single

mode fibers. It is shown that to properly characterize a fiber's bend loss

performance, measurements must be made over a range of bend radii. Further,

that when making measurements at lower bend losses, great care must be taken to

avoid microbend losses. We do not observe the discrepancy seen by others [2,3],

between theoretically predicted behaviour and the experimental results.

Experiment

Bend losses were measured at 1300nm, 1550nm, 1600nm and 1640nm by winding the

fiber on cylinders of varying diameters. The spectral bandwidth was 4nm. The

fiber was laid in a single layer. To avoid microbend losses it was important to

use cylinders with good circularity and smooth surfaces and to wind the fibers

loosely. Experiments were done with fiber lengths as long as 300 m and more

than 1000 turns.

Seven different commercial fibers were investigated. Three were matched clad

(MC) fibers which will be designated A, B and C. Fiber D was by the manufacturer called an MC fiber, but it actually has a slightly depressed cladding.

Fibers E and F were depressed cladding (DC) fibers, and fiber G was a dispersion shifted (DS) fiber.

Results

In Fig. 1 is shown bend losses versus bend diameter at 1550nm for the MC fibers

and the DS fiber. For fibertype A curves for two fibers (Al and A2) are shown.

In addition there is drawn a theoretical loss curve for fiber A based on

Marcuse's theory [4] and where the nominal fiber parameters have been used in

85
the calculations. The losses are given in dB/km, and several tests were done to

confirm that the loss per unit length was independent of the number of turns.

Some of the curves show variations in slope. This is believed caused by whispering gallery modes [5]. The true MC fibers (A, B and C) show on an average no

deviation from the expected theoretical slope as has been observed by others

[2, 3]. Fiber D shows a less steep slope. This is probably due to the slightly

depressed cladding [6].

To achieve good agreement with theory it was necessary to take great care to

prevent loss contributions from microbends . This was especially the case at

larger bend radii where the bend loss was small and the number of turns large.

By using cylinders with round and smooth surfaces and winding the fiber loosely

and evenly spaced, microbend losses could be avoided.

The curve for fiber A2 follows on an average the theoretical curve. Fiber Al

shows a lower loss. Larger variations in loss were seen for other fibers of

type A, but the variations were within the expected limits considering the

specified tolerances on the fiber parameters.

In Fig. 2 is shown bend loss curves at 1550nm for the two depressed clad

fibers. For comparison the bend loss curve for a MC fiber (A2) is also

included. The DC fibers show a marked change in slope with increasing bend

radius. This behaviour is in quantitative agreement with Andreasen's theory for

bend losses in DC fibers [6] . In some regions there is an increase in loss with

decreasing bend radius. Measurements at longer wavelengths indicate that this

effect is caused by whispering gallery modes.

The DC fibers generally show a smaller bend loss than the MC fibers at small

bend radii. At larger bend radii the MC fibers have the lower losses, (Fig. 2).

Discussion

For some applications it is the bend sensitivity at small radii which are of

interest as f.inst. in splice boxes, transmitters and receivers. For cabling

purposes one is concerned with long fiber lengths where bend losses (in the

range of) 0.01 dB/km can be harmfull. In this case it is the bend performance at

larger radii which is of importance [1, 2].

86
The experiments show that for DC fibers the results from small radii cannot be

extrapolated to larger radii. To establish the sensitivity of the DC fibers to

cabling induced losses, measurements must be made at larger bend radii where

great care must be excercised to avoid microbending losses.

For MC fibers it was found that the slope of the loss curves on the average

agreed with the predicted behaviour. But even for these fibers it is necessary

to make measurements over a range of radii to ensure that whispering gallery

modes do not interfere with the results.

It should also be noted that to properly characterize a specific fiber type,

measurements must be made on a number of fibers to establish the spread in bend

sensitivity. Our measurements indicated that the spread varies considerably from

fiber type to fiber type.

Acknowledgement

Optoplan gratefully acknowledges the support of STK - Alcatel, Oslo, Norway for

this work.

References

1. A. Hordvik, M. Eriksrud, "Loss Mechanisms in Single-Mode Fibers Jacketed with

a Loose Jelly-Filled Tube", J. Lightwave Technol. vol LT-4 . pp 1178-1182

(1988)

2. P. Geittner, H. Lydtin, F. Weling, D.U. Wiechert, "Bend loss characteristics

of single mode fibers", 13 tn Europ. Conf. on Optical Comm. (ECOC) , Helsinki,

Finland 13-17 Sept. 1987, Technical Digest, Volume II pp 97-108 (1987)

3. P.K. Bachmann, D. Leers, D.U. Wiechert, "The bending performance of matched

cladding, depressed cladding and dispersion flattened single mode fibres",

0FC/I00C 87, Reno, Nevada, 19-22 Jan. 1987, Technical Digest, Postdeadline

paper PDPl (1987)

4. D. Marcuse, "Curvature loss formula for optical fibers" J. Optical Soc. Am.

66, PP 216-220 (1976)

5. A.J. Harris, P.F. Castle, "Bend loss measurements on high numerical aperture

single-mode fibers as a function of wavelength and bend radius", J. of

Lightw. Techn. LT-4 . pp 34-40 (1986)

6. S.B. Andreasen, "New bending loss formula explaining bends on loss curve",

Electronics Lett. 23, pp 1138-1139 (1987)

87
Bend loss IdB/km)

7 8 9 10 Bend diam. (cm)

Fig. 2

Bend loss vs. bend diameter

for two different DC fibers

(E and F) . The broken line

shows the bend loss for a

MC fiber (A2)

Fig .l

Bend loss vs . bend

diameter for five MC

fibers (Al, A2 , B, C and

D) and one DS fiber (G) . The broken line shows the

calculated bend loss for a fiber of type A using

nominal fiber parameters.

Btnd lost IdB/km)

6 7 8 9

Bend diam. (cm)

88
INDUCED LOSSES IN STRAINED LOOSE TUBE FIBER CABLES

L. Bjerkan, E. Nesset and 0. Aim

EB Cables, P.O. Box 419, N-1371 Asker, NORWAY

INTRODUCTION

In optical fiber cables designed for outdoor use ( aerial, burial

or sea cables ) the most common designs are fibers immersed in a

loose tube or slotted core structure. When such cables are

subjected to high tension or high temperatures, the fibers may be

pressed against the inner wall of the surrounding medium and

microbend attenuation occurs. So far microbend experiments have

been limited to lateral load tests where a short length of fiber

is compressed between plates covered with sandpaper 1 ' 2

, or

compressed against a wire mesh3

, by introducing a short length of

fibers in a vessel filled with small metallic spheres and

applying pressure 4

, or by introducing microbends by a helical

winding arrangement of the fibers where fiber layers cross each

other5 . These experiments have provided interesting data for

microbend attenuation vs. wavelength and fits to existing

theoretical expressions have been achieved4

.

However, whether these experiments adequately describes the

situation in an actual cable is questionable. Therefore new

experiments where the fiber is pressed directly against the wall

of actual tube materials for loose tube cables will be described.

Several fiber types spanning large ranges of production

tolerances of basic fiber parameters have been subjected to these

experiments

EXPERIMENTS

About 140 m of tube was wound in one layer on a large 33 cm

diameter drum. During laying the tube was split in the length

direction using a splitting tool designed for that purpose. This

experimental arrangement allowed for well-controlled and

repeatable fiber insertion under controlled tension. Two tube

materials, polycarbonate (PC) and polybutylene terephthalate

(PBT), which both have been commonly used in loose tube cable

designs have been investigated. The PC tube had inner and outer

diameters of 1.6 and 2.2 mm respectively while the corresponding

figures for the PBT tube were 1.4 and 2.2 mm. Both tubes were

filled with jelly during extrusion, and part of the jelly

remained after splitting and winding the tube. During repetitive

winding an unwinding of fibers, however, the amount of jelly

gradually diminished.

The fibers were wound on the drum one at a time under controlled

tension as shown in figure 1 . During winding the fibers followed

smoothly the bottom of the splitted tube and was in this way

pressed against the tube wall with the tangential tension (cf.

fig. 1) T=mg/2coscc. The relative fiber elongation is then:

e = T/EA = mg / 2EAcoscc ( 1 ) where EA is the effective elastic stiffness of the primary coated

89
fiber which is « 880 N.

The attenuation vs. wavelength was measured using a conventional

spectral attenuation (cut-back) measurement setup. The

attenuation of the fiber configuration on the drum was corrected

for intrinsic fiber attenuation, and the pure strain-induced

losses were found. Three classes of fibers of current single-mode

design were subjected to this treatment: The first was

conventional match clad fibers (MC1 ) designed for 1 . 3 urn operation. The second class was match clad fibers (MC2) designed

for less bending and microbending sensitivity at longer

wavelengths while the third class was depressed inner cladding

(DIC) fibers. All fibers were primary coated with acrylate, but

the coating on the match clad and DIC fibers were of different

types. Fibers have been selected from the available supplies such

that basic fiber parameters mode field diameter (MFD), 2w, and

cut-off wavelength ^ c span wide regions within their specified

limits. Table I shows first the fiber manufacturers tolerances of

these parameters and secondly the actual parameter ranges of the

fibers used in this investigation. Some of the fibers were wound

repetitively on the drums in order to check the reproducibility

of the experiments and also the influence of the diminishing

amount of jelly in the splitted tube.

Table I

General production tolerances and actual range of parameters for three classes of fibers used in this study.

Class of

fiber

MFD and cut-off range

(producers specs.)

MFD and cut-off range

(fibers studied)

MC1 2w=9-11 urn 2w=9 . 82-1 0 . 52 urn *o=1 1 30-1 270 nm * o=1153-1260 nm

MC2 2w=9-1 0 urn 2w=9. 13-9.80 urn ^o=1190-1330 nm 71^=1 21 9-1 320 nm

DIC 2w=8 .1-9.9 urn 2w=8.2-9.2 urn * c=1 200-1 280 nm ^ c =1205-1275 nm

RESULTS AND DISCUSSION

A. Tube materials. The inner walls of the two tube materials were

photographed using a scanning electron microscope. Figure 2 shows

pictures of the two tube materials with the magnification

indicated. The microscopic structure of these materials are

clearly different. Using an image analyzer the maximum amplitude

of the irregularities are «3 urn for the PC tube and ~0 . 5 urn for

the PBT tube.

B. Microbend loss results. Figures 3 and 4 show the range of

induced losses experienced for the three fiber types at e=0.22 %

tensile strain calculated from (1) for the PC and PBT tube

materials respectively. Figure 5 shows an example of induced

losses vs. wavelength with different strains. The results

revealed that the induced losses depend strongly on mode field

diameter and also on the cut-off wavelength of the fibers. Fibers

90
with high values of the MFD and low cut-off were most sensitive

while the opposite condition with low MFD and high cut-off

exhibited the smallest induced losses. This behavior was common

to all three classes of fibers studied. The MC1 class exhibited

the largest sensitivity while the two other fiber designs showed

comparable results. As indicated in fig. 3 the most sensitive

fiber of the MC1 class exhibited considerable losses even at 1.3

urn for e=0.22 % on the PC tube.

The induced losses seem to have a much stronger dependence on

wavelength compared to conventional sandpaper press tests 2

; results also confirmed in our own laboratory. Not surprisingly

the wavelength dependence in our results agrees better with

results from a conventional tension-wind test 2 . However,

experiments with the PBT tube showed less induced losses, about

40 - 60 % at 1.55 um, than exhibited by the PC tube under

otherwise equal experimental conditions. This behavior was

present for all the fibers and is a clear indication that the

microstructure of the tube material contributes to the induced

losses (cf. fig. 2).

By winding the same fiber repeatedly with the same tension at

various stages of the experiments it was observed that the

induced losses gradually became less. This may be explained from

the diminishing amount of jelly in the split tube and that the

jelly itself may possess some kind of microstructure which

contributes to the induced losses. At present this has not been

further analyzed.

CONCLUDING REMARKS

Experiments with strained fibers in a realistic cable structure

have been carried out and induced loss tolerances for current

single-mode fiber designs have been obtained for fibers embedded

in actual tube materials. Efforts are underway to interpret the

results in terms on existing theoretical expressions and to

compare the obtained results with experiments on a strained

complete cable.

ACKNOWLEDGEMENT

The authors thanks the Norwegian Telecommunications Administration (NTA) for partial financial support, and S. Hopland, NTA and

S. Stueflotten, EB Cables for encouraging discussions. We also

thank Corning and Philips for kindly supplying relevant fibers .

REFERENCES

[1] J. A. Dixon, M.S. Giroux, A. R. Isser and R. V. Vandewoestine

Tech. Dig. OFC/lOOC' 87 Reno, NV (1987) paper TuA2

[2] P. F. Glodis, C. H. Gartside, J. S. Nobles: Tech. Dig.

OFC/lOOC 1 87 Reno, NV (1987) paper TuA3

[3] A. O. Garg, C. K. Eoll: Tech. Dig. Symp. on Opt. Fiber.

Measurements, Boulder CO. (1986) p. 125.

[4] M. Artiglia, G. Coppa, P. diVita H. J. Kalinowski and M.

Potenza: Proc. ECOC 87, Helsinki (1987) p. 125.

[5] N. Kamikawa, C.-T. Chang: Tech. Dig. Symp. on Opt. Fiber

Measurements, Boulder CO. (1986) p. 129.

91
Figure 1

Spooling of fiber

in splitted tube

b.

Figure 2

Photograph of inner

tube wall, a: PC,

b: PBT

WAVELENGTH (nm)

Figure 3

Ranges of induced losses

on PC tube, e=0.22 %,

three fiber classes

WAVE L EN GTH (nm)

Figure 4

Ranges of induced losses

on PBT tube, e=0.22 %,

three fiber classes

1250 1350 1458 1550 1650 WAVE LE N GTH ( n m )

Figure 5

Ranges of induced losses

at different strains.

Class MC1 fibers, PC tube

92
MEASUREMENTS OF SINGLE MODE FIBER ATTENUATION

OVER AN EXTENDED SPECTRAL RANGE

Z. Pasturczyk, C. Saravanos, R.S. Lowe,

Northern Telecom Canada Ltd.

Optical Cable Division

Saskatoon, Saskatchewan, Canada

ABSTRACT

A simple modification of the standard cutback method is proposed to allow reliable

measurements of attenuation of single mode fiber for wavelengths shorter than the cutoff

wavelength. Limitations of the standard cutback method were eliminated by optimization of

the fiber layout conditions. Experimental work carried out on depressed cladding and matched

cladding fiber samples demonstrated good repeatability for fiber attenuation measurements:

± 0.020 dB/km at 780 nm, and + 0.010 dB/km at both 1310 nm and 1550 nm.

1.0 INTRODUCTION

Recently, low cost transmission systems for use in subscriber loop networks have been built^

utilizing inexpensive 800 nm lasers and conventional Single Mode Fiber (SMF) optimized for

1300 nm operation. At this wavelength, transmission rates of 150 Mbits/sec over distances of

22.5 km have been achieved by applying special filtering techniques^2

) both to strip higher order

modes and to suppress intermodal dispersion and modal noise. In fact, it has been demonstrated

that when such a system is set up using lasers operating in a single longitudinal mode, it is

attenuation, and not dispersion, limited. Under these conditions, it is imperative to know

precisely and accurately the attenuation of SMF when operating below the cutoff wavelength of

the second, and, even the third, order modes.

Unfortunately, the ordinary cutback measurement technique cannot be used below the factory

measured cutoff wavelength due to the presence of the so-called "mode hump". A mode hump

corresponding to each higher order mode shows up on the traces of attenuation versus

wavelength (see Figure la) if the effective cutoff wavelength of that mode in the "long" test

fiber is considerably different than the corresponding cutoff wavelength in the "short" reference

fiber. In these cases, attenuation measurements within a mode hump are not valid because the

modal content in the "long" fiber is different from that in the "short" fiber.

This paper describes a method for accurate measurements of SMF attenuation parameters over a

wide range of wavelengths by completely eliminating the mode hump effect. This has been

accomplished by modifying the fiber layout conditions such that the fiber cutoff wavelengths

for higher order modes become independent of length.

93
2.0 FIBER LAYOUT OPTIMIZATION

Optimization of the fiber layout involved measurements of cutoff wavelength as a function of

the fiber length and fiber bend conditions (i.e. bend diameter and number of bends). It was

found that the cutoff wavelength dependence on the fiber length can be changed under certain

bend conditions. Reducing the fiber bend diameter attenuates the LPnmode and shifts the

effective cutoff wavelength towards shorter wavelengths. At these shorter wavelengths, the

attenuation of the LPnmode in a straight piece of fiber becomes negligible in relation to its

value at the factory measured cutoff wavelength. Under these circumstances, attenuation due to

fiber bending dominates overall attenuation of the LPnmode and the effective cutoff

wavelength becomes independent of fiber length. Therefore, by taking advantage of this

inverse correlation between the length and bend dependence of cutoff wavelength, the right

combination of bend diameter and number of bends will result in the same cutoff wavelengths

in the "long" and "short" fiber. Once this is accomplished, the mode hump will disappear from

traces of attenuation versus wavelength as shown in Figure lb.

A number of experiments were conducted to characterize the cutoff wavelength dependence on

the fiber bend diameter, number of bends and fiber length. The objective was to determine the

maximum bend diameter and minimum number of bends required to reduce significantly the

length dependence of the fiber cutoff wavelength for two types of single mode fiber; depressed

cladding and matched cladding. Each fiber was approximately 2 km in length.

Cutoff wavelength measurements were done using the multimode fiber reference technique.^

The results are shown in Figures 2, 3 and 4. Figure 2 shows the dependence of cutoff

wavelength (A

c

) on the bend diameter. It is seen that A

c

is a linear function of the reciprocal of

the bend diameter. The slope of each line represents the sensitivity of A

c to bending.

(

4)

Figure 3 shows how the cutoff wavelength (A

c

) changes with the number of bends (N) for a

fixed bend diameter. Note that increasing the number of bends (N) has a similar effect on A

c to

decreasing the bend diameter but since A

c

is inversely proportional to log (N), the dependence

is weak and changing N is not a replacement for changing the bend diameter. It is more like

fine tuning of the optimum fiber layout. Figure 4 shows the cutoff wavelength dependence on

the fiber length for different bend diameres. It is seen that reducing the bend diameter changes

the slope of each line and for bend diameters of 2 - 3 cm, the cutoff wavelength is almost

independent of the fiber length.

It should be pointed out that bending the fiber more than absolutely necessary will attenuate the

LP

Q1 mode at long wavelengths and adversely affect the attenuation measurement accuracy and

repeatability in that region. The optimum fiber layouts selected for fibers used in this work is

shown in Table 1. Fiber attenuation measurements using the layout conditions listed in Table 1

are discussed in the following section.

94
3.0 FIBER ATTENUATION MEASUREMENTS

Fiber attenuation measurements were performed using the standard cutback measurement

technique for single mode fiber. The spectral range of interest was from 750 nm to 1600 nm.

The fiber under test was deployed using the layout conditions shown in Table 1. Both, matched

cladding and depressed cladding SMF samples were measured. Each fiber was measured five

times to check the repeatability of the proposed test method. As Table 1 indicates the achieved

measurement repeatability is + 0.02 dB/km at 780 nm and + 0.01 dB/km at both 1310 nm and

1550 nm.

4.0 CONCLUSIONS

It has been shown in this paper that accurate and precise measurements of SMF attenuation

within the extended spectral range of 750 nm to 1600 nm can be performed using the standard

cutback method if the fiber layout condition is optimized. In addition, it has been observed

that the cutoff wavelength dependence of fiber length A

c(L) is a function of bend diameter.

The tighter the bend the weaker the dependence A

c(L). It has also been demonstrated that the

cutoff wavelength is inversely proportional to the logarithm of number of bends of fiber for all bend diameters down to 1.0 cm.

A final word of caution. The optimum fiber layout introduces a delicate balance between the

attenuation of the LPn

mode and the attenuation of LP

Q1 mode. Different fiber types require

different layout conditions. Depressed cladding fiber needs tighter bends than matched cladding

fiber.

5.0 REFERENCES

[1] M. Stern, W. Way, V. Shah, M. Romeiser, W.C. Young and J.W. Krupsky, in Technical

Digest, Conference on Optical Fibers Communication (Optical Society of America,

Washington, D.C., 1987) paper MD2.

[2] V. Shah, P. Morris, L. Curtis, M. Stern, M.A. Saibi, W.C. Young, in Technical

Digest, Conference on Optical Fibers Communication (Optical Society of America,

Washington, D.C., 1988) paper WQ13.

[3] EIA FOTP-80

[4] L. Wei, R.S. Lowe, C. Saravanos, "Practical Upper Limits to Cutoff Wavelength for

Different Single Mode Fiber Designs". J. Lightwave Technology, Vol. LT-5, No. 9,

p. 1147, 1987.

95
CP

C <-.

c \O CQ

a)

o

u

T3

C

(d

+Jo

>i

(0 (J

s

fi •HP

e

•HW

c

0)6

0)M

w

A3

s;c

o

•H

+J

(0 a 3

o c

S3

a) +j

<

cH

(0 .d

-PCP

c

a) (—

o

>

T3

0) +JO

a> rH

a) w

o

CM

CO<

En

P*» w IN ^

U 1 B

i V N W

fl O +1 o +"l

3

»N

_jm

O -H o +i

NJV ill IN Ml IN ^

in q

C0 H +*l H +|

in a

%r in IN W

U 1 v# rv

CM +1 V ^1 Tl

pft pH UJ (U 1* TJ iH

M pJ C Tl C Tl

• i=1 ft <uw 0J +* r^l _Q g

ft. <rt o

CM CM

• • A •

a) a) 0 TJ a A3 P «

•H >i 0 i—

<G iH

fa H Q U S u

—,9 _

tt cn nj —

(Ui>i/aP) ssoi

ml

ui

A Mn

e

_ B O . i o

g K

s33

o

in oo

o

ino

oo

o

[am] HI HLDN2T2AYM iiOLOO

CQ

* n cm —

S pC

VHC

o e in < H p H p [nu]

3 O O

3 in o

H O O

H H H

MI RLDN3T2AYM ddOUO

8

5

oo

o

o

o

[mi] HI H10N2TJAVM iiOlfD

96
MODE TRANSITION MATRIX TECHNIQUES

FOR MULTIMODE FIBER-OPTIC COMPONENTS

Udo B. Unrau

Institut fur Hochfrequenztechnik der Technischen Universitat

P.O.Box 3329, D-3300 Braunschweig, Germany (FRG)

Abstract

The work on mode transition matrices done by several groups is reviewed and compared.

Applications and practicability of the method are discussed.

1. Introduction

Differential mode attenuation (DMA) and differential mode delay (DMD) cause usually no

trouble in two point fiber links, but DMA becomes important in networks where mode—selective components are concatenated and may cause problems. Since it is impossible to

keep track of a few hundred guided modes Holmes reported in 1981 on the use of a 2*2

mode power matrix [1]. The matrix was measured for two pieces of fiber, and a splice matrix was assumed. He demonstrated that multiplication of the individual matrices of the

fibers and that of a splice predicted quite well the measured response of the concatenated

fibers. Subsequently, this method was elaborated to my knowledge by three groups A-C

around A: R.L. Gallawa (U.S.A.) and J.M. Maisonneuve (U.S.A. and France), B: A.R.

Mickelson (Norway and U.S.A.) and, C: our group. All groups used different approaches to

measure the matrix which will be reviewed in the following. Several names have been

coined for the matrix itself; neither 'transmission matrix' or 'power matrix' should be used

since the first is known in microwaves as a matrix containing full phase information, and

the latter is used to describe simple power transfer between several ports. We prefer 'mode

transition matrix' (MTM) since it is a crude approach to the 'state transition matrix'

(STM) known in electric network theory [1].

97
2. Mode transition matrix measurements

2.1 The mode transition matrix

Holmes [1] has shown that the solution of the coupled power flow equations in a fiber carrying Mmode groups can be expressed in matrix form as [-Pm(^)] = [ffmn(^)] • [-Pn(O)] with m,n = 1...M, where P

p

{z) is the power in the u—th mode group at position z and [#m n(.z)] is the

STM. It is quite impossible to excite a single mode group selectively, therefore, the above

matrix equation has for practical reasons to be simplified to

[P.] = [T.HP] (1)

where [Pi] and [Pj] are the output and input power vectors, respectively, and [Ty] is the

MTM. Most groups used 2x2 matrices, only ours explored indices «,jup to 3. Note that the

length dependence has been omitted in eq. (1) since the MTM concept was soon extended

to characterize fiber components as well [2], and it was pointed out that the real length

dependence in fibers would be extremely difficult to measure due to mode coupling [3].

With ij = 1...3 the MTM is only a very crude approximation to [Hmn] where m,n may span

a few decades.

For use in short—haul networks step—, quasi—step— and graded—index fibers are equally

well suited. Table 1 shows the kind of fiber or fiber components the different groups worked

with to obtain MTM's, and their relevant publications.

Group SI QSI GI

A [4,5] •

B [3,6]

C [7-11,13] [2,9-13]

Table 1: Publications on work with step—index (SI), quasi—step—index (QSI) and

graded—index (GI) fibers and components, done by groups A, B, C

98
2.2 Mode-selective excitation

For measuring a MTM one needs a differential launch — differential detection scheme to

resolve the mode blocks in eq. (1). It is generally accepted that in order to allow easy and

definite mode selective detection, a mode continuum should be excited in a fiber with numerical aperture A^by a source of minimum linewidth AA « A

2

- A^ /

(2tt- a- ni

2

) as shown

in [14]. For A « 820nm and A^x .2 this reads AA » .5nm. We recommend however to

launch a flat mode continuum by a much larger AA [10].

Mode selective excitation is simplest in step—index fibers and can be done by a plane wave

impinging with polar angle #o on the fiber front face. It excites modes with the phase coefficient (3 « ni'k'cos(9o/ni) where n\ is the core refractive index and k the wavenumber. A

spread A$o is directly related to a mode block extending over A/?. This technique was used

by group A [4,5] to obtain 2x2 matrices of pieces of high NA fibers and components. Since

for quasi—step—index fibers (which may have quite odd index profiles) no general optimal

selective excitation exists, the plane wave technique is a good compromise here, too [7].

Several methods have been developed for graded—index fibers. (All groups restricted the

work to parabolic profiles). We preferred gaussian beam excitation through a single—mode

fiber which was transversely offset in front of the GI fiber. A thorough discussion of it together with theoretical and measured selectivity curves can be found in [10]. Group B used a

lens arrangement where the beam can been restricted in its NA and its spot size and was

also transversely moved across a GI fiber endface [3]. A detailed treatment of this excitation method is given in [15]. Unfortunately, the theory used to compute the launch selectivity curves shown in [15] has only been compared to measured near—fields which is quite

uncritical, and since no A—information is given, a comparison with the curves in [10] is

difficult. We noted near—field pattern deformations in the test fiber with the offset method

and discussed it in [10], they were not mentioned by group B.

99
A simple but effective excitation method by plane waves combined with mandrel wrap

filters was reported in [16] with theoretical and measured selectivity curves. In retrospect

we regret that we did not develop it further. Evers found [13] that even without the filter

good selectivity for 2x2 matrices in GI fibers can be achieved.

2.3 Mode selective detection

For step—index fibers, the mode analysis is again very simple and can be directly derived

from the radial far—field, measured either by a detector on a hinging device [4] or by a

vidicon camera. As far as the profile function f(r/a) is monotone, the mode content of all

graded—index fibers can be analyzed with a vidicon camera by differentiation of the radial

near—field—intensity function I(r) with respect to f(r/a) as shown in [17]. Slight corrections

to non—monotone profiles are tolerable as demonstrated on quasi—step—index fibers [10]. It

is mandatory that I(r) is not a function of the azimuthal angle <p and shows no speckle.

Ways to circumvent these difficulties are outlined in [10]. Groups B and C used the nearfield technique for parabolic—index fibers; it proved to be very reliable. Different far—field

techniques exist for graded—index fibers. Rittich derived a useful formula for power—law

profiles [18] and applied it with good success in practical comparison to the proven nearfield technique. We tried a ring aperture technique to resolve the mode block power in

short pulses [12] but found the problems with high precision mechanics nearly unsurmountable. We ran a computer simulation of this modified far—field technique [19] and compared

it with the near—field technique. The latter was superior with less tolerance requirements

and higher stability. The measured field distributions are generally smoothed either by

FFT filtering and curve fitting [6] or by sliding window averaging [10]. A method to reduce

errors in the resulting matrix by checks against other physical requirements and measured

data is found in [13], where also the case of a butt joint between fibers of different radii is

discussed.

100
3. Results, applications, practicability

All groups applied MTM's to DMA- and DMD-concatenation problems, always with good

results, regardless whether 2x2 or 3><3 matrices were used. Matrices were measured for fiber

pieces, couplers, connectors, splices and simple networks built thereof [4—13]. The main

application of the MTM method is aimed at the design of local network structures. In [20]

we tried to assess modal effects in L.A.N. 's with measured matrices as reference. Because of

their high mode selectivity cheap biconically fused couplers are the most critical components in such networks. Provided that otherwise low tolerance components (e.g. fibers with

high centricity, low-loss connectors) are used, we found that no MTM technique is required if up to four couplers are concatenated and if the couplers are characterized by a restricted launch technique. For predicting power level margins in larger networks concatenating such couplers, the MTM method can be very helpful under the same provisions. The

expected power levels and their tolerances at different points of the network should be

quite well predictable. At least, the MTM method is the only known procedure to allow

such predictions. If high tolerance fibers and connectors are used, the MTM method will

fail certainly, too. Comparing the work done so far, 2x2 matrices should be sufficient and

much easier to measure than 3x3 matrices. For all types of fibers the tilted plane wave

excitation, possibly combined with mandrel wrap filters, appears most attractive and inexpensive. At the detection side near—field analysis and Rittich's far—field technique seem

equally well suited. Using the new GaAlAs—Lasers with collimation optics enough power

might be available to allow scanning the far—field with a detector on a simple hinging device. [Pj] in eq. (1) could be measured as the output from a short reference fiber which

could subsequently be butt—jointed to the test object on one of the newly available optimising coupling units. A round robin test on a typical L.A.N, structure would show if the

MTM method is practicable and would help to formulate a standardized test procedure.

101
Holmes, G.T.: Estimation of concatenated system response based on measured transfer functions for

low and high order modes. Proc. 7th Europ. Conf. on Opt. Commun., Copenhagen 1981, paper 3.4

Agarwal, A.K.; Unrau, U.: Characterization of optical components for multimode fiber networks.

Proc. All India Symp. on Fibre Optic Sensors, Bangalore (1983), 39 — 44

Vayshenker, LP.; Hjelme, D.R.; Mickelson, A.R.: Multimode Fiber Systems Characterization. Tech.

Digest Symp. on Optical Fiber Measurements, Boulder, Co. (1986), NBS special publ. 720, 85 — 88

Maisonneuve, J.M.; Gallawa, R.L.: The use of power transfer matrices in predicting system loss: theory and experiment. Fiber and Integr.Opt. 6 (1987), 11 — 26 [based on: Maisonneuve, J.M;

Gallawa, R.L.: Proc. SPIE 500 (1984), 88 - 93]

Maisonneuve, J.M.; Churoux, P.; Gallawa, R.L.: Use of mode transfer matrices in L.A.N, loss evaluation. Proc. SPIE 559 (1985), 182 - 185

Mickelson, A.R.; Eriksrud, M.; Aamlid, S.; Ryen, N.: Role of the fusion splice in the concatenation

problem. J. Lightw. Technol. LT-2 (1984), 126 - 138

Evers, G.; Kober, A.; Unrau, U.: Measurement of mode transition matrices of quasi—step—index optical fiber components. Proc. SPIE 500 (1984), 94 - 99

Evers, G.: Mode transition matrices for fibre—optic connectors. Electr. Lett. 21 (1985), 401 — 402

Unrau, U.: Mode transition matrices as a means for the design of local—area networks with optical

fibers (In German), mikrowellen magazin (Germany) 10 (1984), 644 — 648

Agarwal, A.K.; Evers, G.; Unrau, U.: A measurement set—up for differential mode analysis of fiberoptic components. ntzArchiv (Germany) 8 (1986), 229 — 238

Evers, G.; Unrau, U.: Assessment of modal effects in local area networks. Electron. Lett. 22 (1986),

859 - 861

Evers, G.: Calculation and measurement of mode transition matrices for DMA and DMD characterization of optical fibers. Opt. Eng. 27 (1988), no. 2, 179 — 186 [based on: Evers,G.; Henze, F.; Unrau,

U.: Proc. SPIE 559 (1985), 176 - 181]

Evers, G.: Mode transition matrices for fiber—optic components (In German). Dr.—Ing. thesis,

Techn. Univ. Braunschweig, Faculty of mechanical and electrical engineering, 1987

Mickelson, A.R.; Eriksrud, M.: Mode continuum approximation in optical fibers. Opt. Lett. 7

(1982) , no. 11, 572 - 574

Mickelson, A.R.; Eriksrud, M.: Mode—dependent attenuation in optical fibers. J. Opt. Soc. Am. 73

(1983) , no. 10, 1282 - 1290

Agarwal, A.K.; Evers, G.; Unrau, U.: New and simple method for selective mode group excitation

in graded—index optical fibres. Electron. Lett. 19 (1983), 694 — 695

Leminger, O.G.; Grau, G.K.: Near—field intensity and modal power distribution in multimode graded-index fibres. Electron. Lett. 16 (1980), no. 17, 678 - 679

Rittich, D.: Practicability of determining the modal power distribution by measured near and far

fields. J. Lightw. Technol. LT—3 (1985), no. 3, 652 - 661

Kitayama, K.—L; Tateda, M.; Seikai, S.; Uchida, N.: Determination of mode power distribution in a parabolic—index fiber: theory and application. IEEE J. Quant. Electron. QE—15 (1979), no. 10,

1161 — 1165

Evers, G.; Unrau, U.: Assessment of modal effects in local area networks. Electron. Lett. 22 (1986),

859 - 860

102
A SINGLE LAUNCH TECHNIQUE FOR DETERMINATION OF

MODE TRANSFER MATRICES

by S. Yang, D. R. Hjelme, I. P. Januar, I. P. Vayshenker, and A. R. Mickelson

Department of Electrical & Computer Engineering

University of Colorado

Boulder, Colorado 80309

Introduction:

Component intensive multimode fiber systems such as those presently finding application

in fiber optic local area nets and premise wiring require more characterization than simply loss

statistics. This fact is due to the mode dependency of the fiber compatible components and the fact

that the propagation distances of a component intensive system preclude the achievement of any

steady-state modal distribution. Mode transfer matrices, as were first introduced by Holmes [1] are

powerful tools for multimode system characterization. Unfortunately, as is shown in a recent work

[2], mode transition matrices are launch dependent and, as measurement of a mode transfer matrix

requires various different launches, the mode transfer matrix of a given component cannot be unique.

This lack of uniqueness could be alleviated by standardizing the measurement procedure. However,

a lack of uniqueness tends to put into question the repeatability of a measurement technique.

In the present work, we take a slightly different approach to the standard one. The mode

transfer matrix is actually just a representation of mode transmission function [2], which is an object

which should truly represent the transmission characteristics of a fiber or component. Determination

of the transfer function therefore allows one to determine any of its representations and, therefore, all

the possible non-unique forms of the mode transfer matrices. A simple model of the mode transfer

function for low loss (< 0.5 dB) components is derived and determined experimentally, and the

results of the use of the mode transmission function to derive mode transfer matrices is compared

with direct measurements of mode transfer matrices.

Theory:

The mode transfer function T(R, R') for a component can be defined by the relation [2]

p°(R)m(R) = JdR'T(R,R!)pi (R')m(R') (1)

where p°(R) is the modal power distribution exiting the component, m(R) is the mode density,

p

l (R') is the modal power distribution incident on the component, and R is the mode parameter as

103
defined in any of various works [3]. T(R, R') is a function of two continuous variables R and R! and

is therefore not generally to be found directly from (1). However, one can make a parameterized

model for T{R, R') and use (1) as a fitting equation to determine the parameters. In essence, this

procedure does not differ too greatly from fitting procedures used to determine matrix elements

[4]. To derive a parameterized model, one can assume that the mode-continuum approximation

holds but that coupling only occurs between modes at the same frequency, as was assumed in the

experimentally verified model in reference [5]. The resulting transfer function is

T(R, R') = 6(R - R') - J

dR"m(R")a(R", R)6(R - R') + m(R)a(R, R') (2)

where the coupling function is a characteristic of the component under test. Assuming the component

causes (low) loss and symmetric coupling, one can assume the coupling function to be

«{R,R')=\«Q e-\#-Rn\l* (3)

where ao and r are the fitting parameters.

A Single Launch Measurement Technique:

Once the ao and r of the a(R, R') function are determined, matrices of any order can be

determined from the transfer function. A standard measurement made on components is that of

overfilled loss. The overfilled loss L can be related to ao and r by the relation

L = a0 V

2 r 2r(l + 2r) [l - 2r(l - e" 1 / 27

)] (4)

where V is the fiber V number. Equation (4) can be derived from (1) together with (2) and (3)

where p'iR) is taken to be unity. Clearly, ao can be eliminated from (3) using (4), and the resultant

T(R, R!) is a function of only r. If, when one measures L, however, one were also to record the p°(R)

that results from an overfilled launch, one could use r to fit the relation (1). The resultant best

fitted r could then be used to determine T(R,R'). As this T(R,R') is a complete representation

of the component under test, any size matrix can be generated from it, and therefore the transfer

matrix of the component can be generated from data taken for a single launch.

Experimental Verification:

To truly verify that the measurement technique outlined above is a viable one, one needs

to both verify the ansatz of (3) and to compare matrices generated by the above technique with

matrices generated directly. In order to do this, a fiber concatenation experiment was carried out

104
in which four 1 meter pieces of fiber were spliced together using input and output fibers with five

fusion splices. Comparison of the measured near-fields and those computed using mode transfer

functions are given in Figure 1. Typical comparison of mode transfer matrices calculated from the

mode transfer function are given in Table 1. Agreement in all cases was good.

Conclusion:

By overfilling an input fiber and recording the total loss and near-field output of a test

component, it is possible to find a transfer function for the component. From this transfer function,

it is possible to generate any of the non-unique transfer matrices for this component. This technique

is preferable to presently applied techniques for two reasons. One is simply that only a single launch

is necessary, and therefore intricate submicron scanning techniques of questionable repeatibility are

not necessary. Second, the transfer function contains more information about a component than do

any of its associated transfer matrices. It is safe to conclude that the transfer function approach to

the multimode fiber characterization problem is a most desirable one to follow.

References:

[1] G. T. Holmes, "Estimation of Concatenated System Response Based on Measured Transfer

Functions for Low and High Order Modes," Proc Seventh European Conf on Opt Comm,

Copenhagen, 3.4-1-3.4-4 (September 1981).

[2] S. Yang, I. R Vayshenker, D. R. Hjelme, and A. R. Mickelson, "A Transfer Function

Analysis of Measured Transfer Matrices," manuscript in preparation.

[3] A. R. Mickelson and M. Eriksrud, "Mode-Continuum Approximation in Optical Fibers,"

Opt Lett 7, pp. 572-574 (1982).

[4] See, for example, the technique described in A. R. Mickelson, M. Eriksrud, S. Aamlid, and

N. Ryen, "Role of the Fusion Splice in the Concatenation Problem," Journ Lightwave Tech

LT-2, pp. 126-138 (1984).

[5] A. R. Mickelson, 0. Klevhus, and M. Eriksrud, "Backscatter Readout from Serial Microbending Sensors," Journ Lightwave Tech LT-2, pp. 700-709 (1984).

105
Figure 1: (a) Measured (noisy) and theoretical (smooth) input near-field distribution for overfilled launch, (b) Output near-field distributions measured (noisy) and

calculated from the transfer function (smooth).

L T Tn T12 T2 i T22

.09 .08

Tc .9963 .0066 .003 .9573

Tm .983 0 .001 .978

.16 .085

Tc .9932 .0064 .0052 .9229

Tm .992 .006 .007 .924

.28 .1 Tc .9835 .0115 .0116 .8599

Tm .941 .002 .001 .930

Table 1: Comparison of measured matrices (Tm) and matrices calculated from

the transfer function (Tc) for three splices. The transfer functions are determined

from the same experimental data for the corresponding transfer matrices. L is the

overfilled loss and t is the coupling coefficient.

106
SENSITIVITY OF THE MODE TRANSITION MATRICIES TO

VARIATIONS OF THE INPUT POWER VECTORS

T. J. GRYK, G.E. HOLMBERG, H.M. BLOUNT

NAVAL UNDERWATER SYSTEMS CENTER

NEW LONDON LABORATORY

NEW LONDON, CONNECTICUT 06320

INTRODUCTION:

A matrix method for characterizing the modally dependent losses of fiber optic components was

first reported by Holmes.1 The method has been successfully applied to quasi-step-index

components,2 multimode fiber transmission,3 and graded-index fusion splices.4 The method

uses the near field pattern generated by some type of selective mode excitation to determine

independent power vectors which can be related to the elements of a square transition matrix. In these reports, however, little is said about the exact nature of the independent power vectors,

and to our knowledge, there exists no report on the sensitivity of the matrix elements to

variations of the independent power vectors. It is the purpose of the present work to establish a

relationship between the repeatability of the matrix elements and the independence of the mode

block excitations.

EXPERIMENTAL DETAILS:

The mode transition matrix (MTM) in the present work is a 2x2 matrix requiring two

independent input power vectors to uniquely determine the four matrix elements. The through

and coupled port MTMs of a 3dB, 1 00/1 40um fiber optic coupler were measured for two

different sets of excitation parameters. The measurement set-up is shown in Figure 1 . Variable

mode block excitation is accomplished using a set of apertures which are imaged onto the fiber

endface. Figure 2 illustrates the two excitation states. They are: center/offcenter circular spot

and spot/anulus. Only spatial variations in the launch are made; the numerical aperture of the

beam remaining fixed. It was found that generation of the selective mode blocks yielded a large

group of excited leaky modes. These leaky modes were removed by a mode filter with minimal

mixing of guided modes. The near field pattern for one excitation both before and after mode

filtering is shown in Figure 3. The leaky modes are easily visualized as the "ears" in the near

field. The modal power distribution is calculated from the first derivative of the near field

pattern as described by Daido,5 and then integrated into two independent power vectors. The

division of the mode blocks is arbitrary, and in the present work the division is made at a

relative mode number 5/A=0.5. Four independent power vectors must be measured in order to determine the four matrix elements. They are the input and output power vectors for two

independent excitation states. In order to determine the input power vectors the fiber lead

containing the mode filter was cut back leaving a length of input fiber approximately one meter

long.

EXPERIMENTAL RESULTS AND DISCUSSION:

It is the purpose of this report to discuss the uncertainties involved in a calculation of the MTM

elements. The largest errors introduced into the MTM calculation can be traced to an inability to accurately and precisely measure the power emitted from the optical fiber (3% error is assumed). This uncertainty, however, is mostly random, and therefore the propagation of these

107
uncertainties can be found statistically. The elements of the MTM are determined using

Kramer's Rule:

where

T„ =

P

ol1

?I2

ih1

ih2

det P

MTM =\

Thi =

Tii

lh

P.r P

il1 r

ol1 P

il2 ^12

det P

T,h =

P P

oh1 ih1

P P

oh2 1h2

det P

Thh =

iM oh1

3 P

il2 oh2

det P

and P = ih

?

i, 2

ih 1 Rh2

The notation, although somewhat cumbersome, is required for clarity and is as follows: the

T-values are the matrix elements whose subscripts denote transitions between low(l) and

high(h) order modes [i.e. Thh denotes transitions from high-order to high-order modes]; the

P-values are power levels whose three subscripts denote the exact mode group containing the

power. The first subscript refers to the power as output(o) or input(i), the second as

high-order(h) or low-order(l), and the third as excitation state 1 or 2 [i.e. Pjh2 is the power

contained in the input fiber's high-order mode block when excited by the second launch

condition].

Because the matrix elements are described by the ratio of two determinants, the percent

uncertainty of the elements is found to be a combination of the percent uncertainties of each

determinant. It is convenient to define these as the Normalized Amplitude Factor (NAF) and the

Normalized Independence Factor (NIF), these factors corresponding to the uncertainties of the

determinants of the numerator and the denominator respectively. The percent uncertainty of the

matrix element is then given by:

8T.

M_

TV.

"J

= 2 NIF + NAF

.-2

i|

V

]

PERCENT UNCERTAINTY IN POWER MEASUREMENT

where i,j can have the values of I or h.

and P

M1

Pm2

NIF = -lil^i

P P

112^1

^11^2 + P

il2^h1

NAF., = Pr P

ol1 nh2 Pr P

ol2

r

ih1

Pr P 4. P P

ol1

r ih2 + r

ol2

r

ih1

NAF

h|= P P P P

Tn ^12 _ r\\2 'on

P

H ^12+ P

I2 *ol1 NAF

|h = P P

Th2

Pr P

oh2rih1

^ohlfh2+ l oh2 F?h1

NAF

hh=

P P

Hn rob2 p PTl2

r

oh1

*N1 ?h2 + P

I2 Poh1

The normalized factors defined above have physical significance. The NAF is a measure of the

relative size of the matrix element, while the NIF is a measure of the independence of the input

108
power vectors. The percent uncertainty increases inversely with the magnitude of either the

NAF or NIF. The NAF is a characteristic of the fiber optic component itself. For a component

which weakly couples modes, the off-diagonal matrix terms are small by definition; and

therefore, the percent uncertainty in the measured cross elements will be large. This is seen in

Table I which shows the values of the matrix elements of a 3dB coupler for a given set of input

conditions (fixed NIF). The uncertainty due to the NAF is independent of the measurement

technique, and can only be reduced by increasing the precision of the power measurement. The

NIF is a characteristic of the input power vectors. The NIF becomes small (uncertainty

increases) as the input power vectors become linearly dependent. Therefore, this uncertainty

can be reduced by carefully exciting linearly independent input power vectors. Since the

spot/anulus input power vectors are more independent than the center/offcenter input power

vectors, one would expect the uncertainty to be smaller for the spot/anulus condition. Table I shows this not to be true, however, the uncertainty of the spot/anulus launch being larger than

predicted. [It should be noted, however, that the uncertainty presented with the measured

matrix values is the standard deviation and, as such, only yields a value for the repeatability of

the measurement, not the actual uncertainty.] Therefore, another source of error (possibly

due to leaky mode excitation or near field irregularities) limits the precision of our measurement and is currently under investigation. As the NIF becomes very small, this other

source of error becomes insignificant, and the theoretical uncertainty matches the measured

uncertainty as shown by a comparison of the measured and expected uncertainties for the

center/offcenter excitation found in Table I. Another key feature of the matrix elements is their

invariance to the specific set of independent input vectors. The discrepancy between the mean

values of the matrix elements calculated with the above input excitations is insignificant. Only

the uncertainty is a function of the vectors, and therefore any set can be used to determine the

matrix elements. It should be noted, however, that if strongly dependent input conditions are used, the precision of the measurement will become vanishingly small.

In conclusion, we have measured the repeatability of the matrix elements as a function of the

independence of the input power vectors, and compared them to the uncertainty predicted from

statistics. We have found that the values of the matrix elements are independent of the

variations of the input power vectors, however, the uncertainty of the measurement varies

inversely with the magnitude of the Normalized Independence Factor.

REFERENCES:

1 Holmes, G.T., "Estimation of Concatenated System Response Based on Measured Transfer

Functions for Low and High Order Modes," Proc. 7th European Conference on Optical

Communication, Copenhagen, 1981.

2

Evers, G., Kober, A., and Unrau, U., "Measurement of Mode Transition Matrices of

Quasi-Step- Index Optical Fiber Components," SPIE Vol. 500 Fiber Optics: Short-Haul and

Long-Haul Measurements and Applications II, 1984.

3Vayshenker, LP., Hjelme, D.R., and Mickelson, A.R., "Multimode Fiber Systems

Characterization," Technical Digest: Symp. Opt. Fiber Measurement, NBS Special Publication

720, 1986.

4

Eriksud, M., Mickelson, A.R., Aamlid, S., and Espe, B., "Mode Dependence of Splice Loss in Graded-lndex Optical Fibers," IEEE Jour. Quant. Elec, Vol. QE-19, No. 5, 1983

5Daido, Y., Miyauchi, E., Iwama, T., and Otsuka, T., "Determination of Modal Power Distribution

in Graded-lndex Optical Waveguides From Near-Field Patterns and its Application to

Differential Mode Attenuation Measurement," Applied Optics, Vol. 18, No. 13, 1979, p.2207.

109
A

CHROMATIC

FILTER

VIEWING CAMERA

MODE

> 1 0X FILTER

MICROSCOPE

OBJECTIVE

0

in

VIDICON

IMAGING

SYSTEM

WHITE

LIGHT

SOURCE

VARIABLE

APERTURE

;TEST FIBER

THREE AXIS

LAUNCHER l DIGITIZER

DESKTOP

COMPUTER

FIGURE 1 . APPARATUS USED TO DETERMINE THE NEAR FIELD

OF A 100/1 40 yn GRADED-INDEX FIBER

FIGURE 2. SELECTIVE MODE

EXCITATION

leaky modes

i

0.0 1.0

NORMALIZED RADIUS

1.0

2

o.o

no leaky modes

0.0 1.0

NORMALIZED RADIUS

FIGURE 3. NEAR FIELD PATTERN BEFORE

AND AFTER MODE FILTER

CORE/ANULUS EXCITATION

THROUGH PORT COUPLED PORT

.59 +.08 .01 + .06 \ measured I .36 ±.03 .07 ±.04

.00 +.02 .43 +.07 J

( 7 tria,S)

I .05 +.01 .36 +.02

.05

.05

.11

.04

calculated

uncertainty

.04

.05

.01

.03

CENTER/OFFCENTER EXCITATION

THROUGH PORT

.58 +.08

.00 +.03

.08

.01

.04 +.06

.39 +.06

.08

.05

measured

(7 trials)

calculated

uncertainty!

COUPLED PORT

.31 +.01 .24 + . 09

.06 +.02 .35 +.06

.07

.06

.04

.05

TABLE I. MATRIX VALUES WITH MEASURED AND CACLULATED

UNCERTAINTIES FOR VARIOUS INPUT CONDITIONS

110
EVOLUTION OF POWER DISTRIBUTIONS IN FIBER OPTIC SYSTEMS

- DEVELOPMENT OF A MEASUREMENT STRATEGY

M. J. Hacker

Corning Glass Works

Corning, N. Y. 14831

Introduction

System power budgeting using multimode fiber is made more complex by the

dependence of the actual loss experienced in the system on the power

distribution carried by the components. Also, each component comprising the

system has the potential to alter the mode power distribution (MPD) carried

by the fiber . Thus, the actual loss introduced by each component is

controlled by the components which precede it in the system, as well as the

power distribution launched by the source. This paper presents a basic

description of MPD evolution through components and some experimental results

of component characterizations which illustrate these behaviors.

Equilibrium mode distribution (EMD) and steady state MPD are commonly used

interchangeably. However, they have very distinct definitions . An

equilibrium MPD has the property that irrespective of the input power

distribution, the output distribution is consistently the same as some MPD

which is characteristic for the device under test. In a steady state MPD, on

the other hand, the input and output MPD are the same. Thus, a steady state

MPD is not necessarily defined by the system element, but rather is dictated by

the launch condition.

Fiber Characterization

The propagation of power through a fiber is relatively easy to understand and

characterize since high quality fibers generally carry steady states - the

output MPD is equivalent to the input MPD. However, if the launched

distribution puts power into the higher order modes of the fiber, which are

loosely bound or do not propagate to the end, the fiber will act as a mode

filter. Using a uniformly overfilling source, MPDs at the end of a length of

fiber clearly show this filtering as a reduction in higher order mode relative

power. Also, the mode filtering occurs gradually over some distance which is

evident when the transient loss is plotted versus length.

A number of steady states can exist for fiber. As long as power is not

launched into the transient modes, the fiber will carry the steady state.

Examples of common power distributions (OFL, LPS, and an ultra-restrictive

launch) are given in Figure 1. The existance of a steady state is confirmed by

noting that the 2 m power distribution is equivalent to the 2.2 km MPD within

the measurement reproducibility except for the OFL case, where mode filtering

would occur.

Ill
Component Characterization

In general, components are much more difficult to characterize because they

can produce an equilibrium MPD as opposed to a steady state. An equilibrium

producing component will filter the MPD in a similar fashion to a steady state

producing component if the launch power distribution exceeds the equilibrium

MPD. If, on the other hand, the launch MPD is more restrictive than the

equilibrium MPD, the component will mix power into some higher order modes

to produce the equilibrium MPD.

An important similarity between fiber and components is that it generally takes

a few hundred meters of fiber or a couple of components to develop a steady

state or equilibrium, respectively. Therefore, in order to study the

cumulative effect of a particular type of a component in a system, a number of

the components (6 to 10) need to be joined in series with low-loss fusion

splices.

As an example, 6 wavelength^ division multiplexers / demultiplexers fabricated

by ion exchange technology ' ' were concatenated. The end-to-end loss for

three different launch conditions - OFL, LPS, and an ultra-restricted launch

(15 im / 0.1 launch NA) - was measured as each component was removed from the

concatenation, and the results plotted in Figure 2. The sum of the individual

component measurements for each launch condition is also included. For the

OFL and ultra-restricted launches, end-to-end and sum loss measurements diverge

rapidly for increasing component count. Thus, OFL overestimates the actual

loss experienced (by as much as 6 dB), and the ultra-restricted launch

underestimates it. It can then be concluded that, because the error increases

with increasing component count, neither of these launch conditions are

adequate for prediction of actual system loss.

It can be noted in Figure 2 that the slopes of the concatenation losses tend

toward the same value for increasing component count in the case of the LPS

measurement condition. This suggests that, just like with fiber, there is a

steady state attenuation (approximated in this case by LPS) combined with a

transient as equilibrium is achieved. To clarify this point, the loss added to

the concatenation for each launch condition as each component was added was

calculated. The difference in loss between this added loss and the LPS loss

measured on the component is given in Figure 3. Because this difference in

added loss when an LPS launch is used most closely approximates zero, the LPS

power distribution most closely approximates the equilibrium power

distribution. This result is convenient. However, it is not necessarily true

that other types of components will have the same ability to use LPS to

approximate their equilibrium loss.

There are two ways to verify from our data that equilibrium was indeed

developed in these particular components. First, equilibrium is implied by the

ultra-restricted launch loss. In Figure 3, the difference in ultra-restricted

launch loss shows a transient for the first few components (similar to OFL but

with a different sign), but goes to zero. This implies that the low order mode

power of the ultra-restricted launch loss is being mixed up to higher order

modes. The resultant power distribution produces a higher loss and is

characteristic of the equilibrium.

112
The second way to confirm that equilibrium is actually being developed is to

measure the MPD for an overfilling and ultra-restricted launch. If the exiting

MPD is the same at the end of the concatenation, the power distribution is that

of an equilibrium by definition. The MPDs in Figure 4 show this to be the

case.

Conclusions

From the results presented here, fibers and fiber-optic components can be

classified as either steady state producing filters or equilibrium producing

mixers - typically fiber or components, respectively. By assuming that the

system will be comprised of filters and/or mixers, accurate system design rules

can be greatly simplified. The work to date suggests that the steady state

loss combined with a one time transient due to mode mixing and filtering can

optimally characterize a component. By doing so, the double book-keeping of

the transient when a non-steady state loss is used in the prediction of a

system (a concatenation of components) should be eliminated. However, further

work in the form of experimentation and model system testing is necessary to

develop a simple set of design rules.

References

(1) M. Calzavara, P. DiVita, and U. Rossi, "Mode Power Distribution

Measurements in Cptical Fibers", CSELT Rapporti Tecnici , Vol. XI,

no. 5, p. 447-451, October 1981.

(2) P. Reitz, M. Hackert, J. Matthews, K. Murphy, D. Hanson, M. Hartmann, T.

Huegericn, K. Kevern, T. Odderstol, "A Comparison of Loss Measurement

Methods with Observed System Loss in Optical LAN Links", presented at the

SPIE Fiber Optics Technical Symposium - 559-24, San Diego, California,

August 21-22, 1985.

(3) M. Hackert, "Development of Equilibrium Mode Distribution (EMD) in a

Fiber", presented at the Second International SPIE Technical Symposium -

584-28, Cannes, France, November 25-29, 1985.

(4) G. Chartier, P. Callier, A. Guex, P. Faussaucl, and Y. Won, "Graded-Index

Surface or Buried Waveguides by Ion Exchange in Glass," Applied Optics , Vol. 19, pp. 1092-1095, April/June 1980.

(5) A. Beguin, T. Dumas, M. Hackert, R. Jansen, C. Nissim, "Fabrication and

Performance of Low Loss Optical Components Made by Ion Exchange in Glass,"

presented at Optical Fiber Conference '88, New Orleans, Louisiana, January

25-28, 1988.

(6) K. Murphy, "Integrated Optics on Glass," Photonics Spectra , November, 1987.

113
50/125/0.2 Fiber

Mode Power Distributions

Figure 1

Cumulative Loss Performance

WDM Concatenation

Figure 2

Number Of WDM Devices Relative Mode Number

Equilibrium Loss Development

In A WDM Concatenation

Figure 3

Mode Power Distributions Of A

Component Concatenation

Figure 4

114
Accurate Measurement of Optical Power in Fibre Optic Systems

D H Nettleton

National Physical Laboratory Teddington, Middlesex TW11 OLW, UK

INTRODUCTION

Optical power is measured in fibre optic systems with commercially

produced fibre optic power meters fitted with an appropriate connector.

If power measurements are to be made accurately these meters must be

calibrated against national spectral responsivity scales. Traditional

calibration facilities are not suitable to carry out these calibrations

because of the difficulty in coupling radiation into the meters except

with an optical fibre and so special facilities and transfer standards

have to be developed.

ABSOLUTE RESPONSIVITY

Absolute responsivity calibrations of fibre optic power meters are

carried out at NPL against a specially developed transfer standard

detector. This is shown in Figure 1. It is constructed with a 5 mm

diameter germanium photodiode mounted on the exit port of a 50 mm

diameter integrating sphere. An FC-type connector is mounted on the

input port. The response of this detector is measured against calibrated

thermopile radiometers using laser diode sources at 8^0 nm, 1300 nm and

1550 nm. Radiation is imaged from the sources into the integrating

sphere or onto the calibrated thermopile radiometer in an automated

calibration rig. The integrating sphere removes the sensitivity of the

response of the detector to image size and position. This eliminates

uncertainties caused by non-uniformities in the response across the

surface of the germanium photodiode. It also makes it possible to change

the connector type at the input port without altering the response of

the detector. Finally reflections from the germanium photodiode have a

negligible affect on the system to be measured. The only disadvantage of

this transfer standard is the reduction of the signal by three orders of

magnitude. However, if calibrations are carried out at a power level of

100 uW {- 10 dBm) , uncertainties can be kept very low. Calibration using

single mode fibre has been achieved with an uncertainty of + 0.25%

(+ 0.01 dB) and with multimode fibre + 0.5% (+ 0.02 dB) . Replacing the

115
germanium photodiodes with silicon photodiodes for measurements at

850 nm and InGaAs photodiodes for measurements at 1300 nm and 1550 nm

will reduce the noise level and make possible measurements at lower

power levels.

LINEARITY

The response of fibre optic power meters at other power levels can be

measured using a cooled 5 mm diameter germanium photodiode fitted with a

suitable connector cap. If all measurements are made relative to the

100 uW (- 10 dBm) absolute response calibration, the long term stability

of the cooled detector is not important; nor is the patch size and

position as long as it is maintained throughout one set of measurements.

RELATIVE SPECTRAL RESPONSE

The change in response of fibre optic power meters away from the spot

wavelength calibration of the laser diodes is carried out by reference

to 5 mm diameter germanium photodiodes which hold the relative spectral

response scale at NPL. This scale has been established with pyroelectric

detectors which have a platinum black surface behind a silica window.

The relative spectral reflectance is calculate from the measured

reflectance of the black and the transmission of the window. The

uncertainty of the scale is shown in Figure 2.

Fibre optic power meters are calibrated by recording the response to

monochromatic radiation from a spectrally dispersed incandescent source

imaged either onto the sensing element with the connector cap removed or

into a fibre connected to the meter. Measurements are carried out at

20 nm intervals and the relative spectral power falling on the power

meter measured by reference to the germanium photodiode standards.

ERRORS DUE TO INTER-REFLECTIONS

Solid state photodiode detectors can have high reflectivities even when

anti-reflection coated. Typical reflectances are between 20% and h0%.

Care must be taken, therefore, when connector caps are fitted as they

may significantly increase the response of a power meter by

inter-reflections. Increases as large as 16% (0.6^ dB) have been

measured. Of course if the meter is calibrated with the cap in place the

116
correct response will be recorded. However the calibration may not hold

if different caps for different detector types are used. Because

anti-reflection coatings are wavelength dependent, the Increase in

response due to inter-reflections may also be wavelength dependent.

White diffusing paint

Detector

Fibre optic

connector

Integrating

sphere

Figure 1. Transfer standard for the calibration of fibre optic power meters

rmTTTl TTTTTTT T TTT I I I I I I I I I T Trm rTTTTTI ITT T TTT I TTTT I I I I I I 1 T mt ITTI TTTT n"n TTTT IT IT

C

o

u

in . CD I

c

ou

k.

4/ 0.

0 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 )i 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 . 1 1 1 1 1 1 1 1 1 i eoo 1000 1200 1400

Wavelenglh/nm

1600 1800

Figure 2. Uncertainty in NPL relative spectral responsivity scale

117
118
An Accurate Method of Determining Single-Mode Fiber Cutoff

Wavelength

C. N. Okafor

D. W. Peckham

AT&T Bell Laboratories

2000 Northeast Expressway, Norcross, GA. 30071

1. INTRODUCTION

An important parameter for a single- mode optical fiber is the effective cutoff wavelength,

because it indicates the minimum wavelength at which a system can operate efficiently. The

term effective cutoff wavelength is defined as the wavelength where the first higher order mode

(LPj*) starts to propagate and increases the total power transmitted through a fiber (of a given

length and configuration) by 0.1 dB.

The mode suppression technique is the most common technique used to measure the effective

fiber cutoff wavelength (EIA-455-80). The power through a short length of fiber (usually 2m) is measured at selected wavelengths (typically every 10 nm) both with and without a small

diameter mode suppression loop. The cutoff wavelength is the wavelength at which the "loop

loss" is 0.1 dB. When the 0.1 dB loss value falls between measured data points, linear

interpolation is used to determine the fiber cutoff wavelength. This estimated cutoff wavelength

tends to be higher than the actual fiber cutoff wavelength because the power vs. wavelength

curve near X

c

is concave rather than linear (Figures 1 and 2). The measurement accuracy can be improved, at the expense of increase of measurement time, by decreasing the spacing

between measured wavelengths.

In this paper, a more accurate method of determining the fiber cutoff wavelength by fitting an

emperical model to the loss data is introduced. The advantages of this method over linear

interpolation are:

1. Fewer data points, spaced 20 - 40 nm apart, can be used in the curve fit to achieve a given

measurement resolution.

2. Fewer data points mean shorter measurement time, and a resulting savings in cost.

3. The determined cutoff wavelength is more accurate because it accounts for the actual

shape of the "loop loss" vs. wavelength curve.

4. The precision is increased because a least-squares curve fitting procedure is used.

2. DERIVING AND LINEARIZING THE LOOP LOSS EQUA TIONS.

An expression for the cutoff wavelength measurement "loop loss" is derived that more

accurately describes the LP^^ mode behavior near its cutoff than the straight line interpolation

used in the EIA procedure.

We assume that a) the first higher order mode (LPjj) is so weakly guided near cutoff that it is completely attenuated by the loop inserted during the fiber cutoff wavelength measurement,

and b) the attenuation of the fundamental mode (LPqj) caused by the loop and short length of

fiber is negligible. Both of these conditions are generally met near X

c

if the diameter of the loop

is properly chosen (typically 25-50 mm). Therefore, only the LPqj mode power Pq- exits the

fiber when the loop is present, and P«, + P, , 10-o,n

L/ 10 when the loop is removed.

Ql + pu

10-0,111710 when tne loop is removed. Pni an< Pjj are the initial launched powers guided uded by the LPqj and LPjj modes, respectively, respectively, or^ is the LPj^ mode attenuation in dB/m, and L is the fiber length in meters.

Thus, the loss of the loop is written as:

Loss . 10,og

(P,. + Pnl0-°"^°)

119
Now assume that the power launched into the LPjj mode is a multiple factor of the LPgj

modal power represented by p, (p « 2), so that P^j = pPqjThe wavelength dependent attenuation of the LPjj mode, can be written as [1,2]

AX

* X ~ x

c>/ 10

11 (2)

where A = lOlog

^q0 oi ~ (A=19.3 for p=2), x is a constant related to the fiber index profilef

(x ^ 1.4 for AT&T depressed cladding fibers), X is the wavelength where a^* is evaluated, X

c

is the cutoff wavelength to be determined, and L is the length of fiber at which X

c

is defined

(generally 2m).

Substituting Eq. 2 into 1, with P^j = ^Pq^ gives

Loss = lOlog 1 + plO

_Ax(x - x

c)/io

10 (3)

We can simplify Eq. 3, the "loop loss" equation, to the form of a linear equation in X by

jqLoss/10

j rearranging and letting Y = lOlogi —7~l°g . Eq. 3 becomes:

Y = (X - X

c )logX (4)

This is a straight line equation in the X and Y plane, i.e, Y = a + bX where a = — X

clogx is the

Y-intercept and b = logx is the slope.

Eq. 4 has two unknown parameters X

c and x which can be uniquely determined by using the

same two data points as in the EIA linear interpolation procedure. (With additional data points, a more robust regression analysis can be performed as discussed in section 3.)

To compare the cutoff wavelengths determined by the two methods, a fiber was measured every

2 nm in the region around the cutoff wavelength. The loop loss model (Eq. 4) was fit to this

data to accurately determine the wavelength where the "loop loss" is 0.1 dB, i.e, X

c

- Next, the

cutoff wavelength using the EIA linear interpolation procedure, X

EIA

, was determined using

data points spaced 10 nm apart. Using the same two data points tnat were used to determine

X

EIA

, the loop loss model was used to determine X

c

2pts

. Figure 2 shows the plots of the results

of these 3 calculations. The data points in the square boxes are the ones used to determine

X

c

EIA and X2pts

. Figure 2 shows that X

2pts

is very close to X

c because the model accurately

represents the loss curve, however, the EIA method using linear interpolation, results in an

overestimate of X . c

S. DISCUSSIONS OF THE ROBUST CUTOFF WAVELENGTH

The loop loss model can also be used for a more robust estimate of cutoff wavelength by using a

linear least-squares regression analysis in the X and Y plane to fit the model to more than two

data points. The linear regression analysis gives the slope b and the intercept a of the linear

equation of which X

fi = — a/b and x = 10b

-

f

It can be shown x > s related to the length scaling of the cutoff wavelength. For a straight fiber, the cutoff

wavelength of a fiber of length L is X„(L) - \A2) - T^log(—), where X (2) is the 2-meter cutoff wavelength.

c c logx *

120
When two data points are used, the model yields a cutoff wavelength, X

c

pts

, that is more

accurate than the linear interpolation method used in the EIA procedure. However, both

methods force their respective interpolating functions through two particular data points, and

the results are sensitive to any error in the measurements. Performing a linear least-squares

regression analysis with more than two data points, spaced by 20 nm or more, has the

advantage of:

1. Averaging the measurement uncertainties and yielding a more precise value of X

c

.

2. Including additional data points below the cutoff wavelength where potential modal noise

generation is a concern.

In a fiber system, some fibers may be operated below their cutoff wavelength, and this causes concern about modal noise generation. Including two or more points within this region in

determining the X

c of the fiber yields a better representation of the LP*, mode attenuation in

this region of interest. Therefore, the determined cutoff wavelength will be a more accurate

estimate (considering modal noise) when specifying fibers.

Figure 3 shows the linear regression analysis on X and Y, and a plot of X and loop loss (dB) for a

depressed-cladding single-mode fiber. The data points are spaced 20 nm apart and those in the

square boxes are the ones used in the regression analysis 4 Also plotted are three vertical lines

showing the cutoff wavelengths determined by using 1) linear interpolation, X

c

LIT

, 2) the model

with the same two data points used in the linear interpolation X

c

2pts

, and 3) the model with 3 or more data points, X

cfit

. The cutoff wavelengths estimated by the three methods are 1267.4,

1263.6 and 1264.3 nm, respectively.

Table 1 shows tabulations of the calculated cutoff wavelengths using three different methods

(X

cEIA

, X

2pts

, X

c

cfit ) compared to the "actual" cutoff wavelength, X

c

, determined by measuring

additional data points spaced 2 nm apart in the cutoff wavelength region and fitting the model

through the points. The data points used in determining X

c

cfit and X

c

2ptg are spaced 20 nm

apart while the ones used in determining X

cEIA are spaced 10 nm apart. The average error of

the EIA method is higher than the other two methods where the model is used, even though the

data points are spaced 20 nm apart.

4. CONCLUSIONS

The model described in this paper determines cutoff wavelength more accurately and with fewer

measurements (and hence, lower cost) than by using the standard EIA-455-80 procedure.

Fitting two data points with the model gives a more accurate estimate of the "traditional" fiber

cutoff wavelength (even if the data points are spaced 20 nm apart) when compared with the EIA

method. Including 3 or more loss values in the 3.0 - 0.03 dB range gives a robust value of cutoff

wavelength that incorporates the concern of potential modal noise generation. Using the model,

20 to 40 nm spaced data points can be used with excellent accuracy. This new model works well

with both depressed and matched cladding fiber designs.

REFERENCES

[1] D. G. Duff, F. T. Stone, and J. Wu, "Measurements of Modal Noise in Single-Mode

Lightwave System," Technical Digest, 8th Conference on Optical Fiber Communication,

San Diego, California, (Feb. 11-13, 1985), paper TU01.

[2] F. T. Stone, "Modal Noise in Single-Mode Fiber Communications System," Proceedings of

the SPIE, San Diego California, vol. 500, pp. 17-22, (1984).

t "Loop loss" values in the 3.0 - 0.03 dB range are used in the regression analysis. At higher loss values some of the

assumptions of the model are no longer valid, and below 0.03 dB small measurement uncertainties can dominate.

121
Table 1. Comparism of the cutoff wavelength determined by the three different methods

•with the actual cutoff wavelength.

Fiber Cutoff Wavelengths

Fiber ID #

Errors

Actual X

c X EU - X

c c X

2plt - X

c c x,

cnt - X

c c Fiber # 1 1294.2 nm + 1.6 nm -0.2 nm + 1.0 nm

Fiber #2 1309.3 nm +0.6 nm +0.5 nm +0.3 nm

Fiber #3 1335.7 nm +0.8 nm -0.1 nm -10 nm

Fiber #4 1331.6 nm + 1.6 nm + 1.2 nm +0.2 nm

Average Error + 1.15 nm +0.35 nm +0.12 nm

0.25 -

0.20

0.15

408

0.10 |-

0.05

0.0 fc_

0

•o

ec

iu1 5.0

o,a

Ui| 3.0

toz

< 2.0

> LENGTH OF FIBER f WITHOUT LOOP

1.0 -

00

LENGTH OF FIBER

WITH A LOOP

FIBER CUT-OFF

WAVELENGTH

0.9 1.0 1.1 1.2 13

WAVELENGTH fam)

1.4 1.5

FIG. 1. Normalized Plot of Transmitted Power

With and Without a Mode Suppression Loop

4.0

1294.16

X?* 1295.79

. X

2P«» . 1293.97

- 2

-10

-12

3 -14

1285 1290 1295 1300

WAVELENGTH (nm)

1305 1310

FIG. 2. Plots of X , estimated by

three different methods.

1180 1180 1200 1220 1240 1280 1280 1300 1320 1340 1380

WAVELENGTH (nm)

FIG. 3. Plots of X and Y, X and

Loop Loss (dB) , and the estimated

X .

122
A Technique to Estimate the Cutoff Wavelength Profile in

Single Mode Fibers Using a Switchable Dual Wavelength OTDR

by

K.W. Kowaliuk and J. Ferner

Northern Telecom Canada Limited

Optical Cable Division

Saskatoon, Saskatchewan, Canada

ABSTRACT

A technique is presented which provides a measurement of the cutoff wavelength profile in

single mode fibers using a dual wavelength OTDR. The theory of this technique is based on

the dependance of the backscatter capture fraction on the mode field diameter and on the

dependance of the mode field diameter on wavelength and cutoff wavelength. Accuracies of

± 20 nm in the measurement of the relative deviation of A

c are reported for typical single mode

fibers. However, in certain circumstances caution must be exercised when interpreting results.

BACKGROUND

The backscatter signal obtained from the OTDR measurement of an optical fiber is the result of

the combination of three key mechanisms; power decay or attenuation, local scattering and the

backward capture and transmission of the scattered light. In its simplest form, the backscatter

power P received from a given point z in a fiber can be represented as the product of these

effects.

P(z) a P

Q

a (z) B(z) exp[ -2 f^Wdx ] (1)

J o

where P is the power launched into the fiber, a (z) is the local scattering coefficient,

B(z) is the backscatter capture fraction, and i(z) is the local attenuation coefficient.

A technique developed by DiVita and Rossi [1] provides a highly reliable means of separating

the decay and waveguide imperfection effects from an OTDR signature. For OTDR signatures

S

x(z) and S2(z) (which include the processing performed in the OTDR) launched from opposite

ends of the same fiber of length L, the imperfection contribution I(z) becomes:

I(z) = S^z) + S

2(L-z) = 10 log

10 [a (z) B(z)] + Ax (2)

while the attenuation (decay) contribution D(z) is:

D(z) =

S

i(z)

I S

2(L~ Z) = A2 - A

3 f

e7(x)dx (3)

2

J o

where Ar

A, and A

3 are constants. The above expressions are derived assuming a pulse of

negligible width. For practical applications, the finite width of the laser pulse must be

accounted for by shifting the signatures relative to each other before the separation process.

Correct interpretation of the imperfection line requires an understanding of the mechanisms

involved. For a typical single mode fiber with a Gaussian field distribution, it has been shown

[2] that the backscatter capture fraction B(z) is proportional to (A/w)2 where A is the wavelength

of incident light and w is the LP

Q1 mode field diameter. It can also be shown that the influence

of the local scattering coefficient as(z) on the imperfection line can be neglected compared to

the MFD contribution.

123
An empirical expression was developed by Marcuse [3] which relates the mode field diameter of

a step index fiber at a given wavelength to the fiber V number and core diameter. The

expression can be rewritten to be a function of wavelength, LPn

cutoff wavelength and core diameter. To eliminate the core diameter dependance, the expression is evaluated at two

wavelengths Aj and A

2 and the two expressions divided. The ratio of mode field diameters, R,

can be written as:

0.65 + 0.434 + 0.0149 $)°

R= = ^— —

6

(4)

0.65 + 0.434 j^j + 0.0149

w(A,)

For simplicity of processing, equation (4) was converted to a more solvable form. For test wavelengths in each of the 1310 nm and 1550 nm regions, equation (4) can be approximated

(over the cutoff wavelength range of 1100 nm to 1400 nm) by a second order polynomial of the

form:

A

c = C1R2 + C2

R + C

3

< 5>

where cr

c

2 and c3 are constants.

ANALYSIS ALGORITHM

After obtaining and separating the signatures according to equations (2) and (3), the

imperfection lines at each wavelength are referenced to the arbitrary reference point z= z

q

. The

modified imperfection line I'(z) can then be represented as follows:

I'(z) = I(z) - I(z

o

) = 20 log

10

(-JSgjl

]

(6)

By taking the difference between the modified imperfection lines at both wavelengths, the

expression becomes:

AI(z) =I-(A

2

, z) - lXXv

z) = 20 log

l0

• -^1) (7)

Expressing equation (7) in terms of the MFD ratio R gives:

R(z) = R(z

o

)

• 10 ( AlW/2°) (8)

R(z

q

) can be determined either by estimating or physically measuring the cutoff wavelength

A

c(zQ

) and substituting this into equation (4). R(z) is then calculated from equation (8) which is then used in equation (5) to determine A

c(z). Although this calculation will produce an absolute

profile of A (z) based on the estimated value of A (z

Q

), it is preferable to present the result of

this technique as a profile of the deviation in cutoff wavelength rather than an absolute cutoff

wavelength profile. Figure 1 illustrates the various stages of this analysis technique.

EXPERIMENTAL

To verify the reliability of this technique, a number of fibers were produced with severe

variations in cutoff wavelength along their length. All the fibers produced for this study were

dispersion-unshifted-depressed-clad single mode fibers. OTDR signatures were obtained from

each fiber and were separated and analyzed using the technique described above. Each fiber

was then cut at regular length intervals and a sample of fiber was taken from each position.

The samples were then measured for cutoff wavelength according to EIA-455-80.

124
Test results from one of these fibers (labelled fiber A) is shown in Figure 1(e). From the test results produced, the technique has proven capable of predicting relative cutoff wavelength

variations with an accuracy of ± 20 nm, provided the appropriate pulse width is used.

LIMITATIONS OF TECHNIQUE

There are several key assumptions made in the development of the analysis algorithm which

must be taken into account when interpreting the results of this technique. Some of these

assumptions include step index behavior of the fibers and the Gaussian form of the mode field

diameter. In addition, the effects of the digital filtering performed internally to the OTDR

equipment have been neglected. Indeed, it would be possible to develop a more detailed theory

to take these factors into account. However, based on the results achieved to date, this

technique is limited more by the resolution and signal-to-noise of the reflectometer than by the

theory itself.

Nevertheless, one key limitation to this technique is that the theory used assumes that the cutoff

behavior of the LPU

mode can be predicted based on the behavior of the LP

Q1

mode. Provided

the shape of the refractive index profile is known, the relationship between these modes can be

predicted. However, in order for the prediction to be effective, the general shape of the

refractive index profile must remain relatively consistent throughout the fiber length so as not

to affect the behavior of one mode relative to the other. This can become particularly

important when analyzing fibers with refractive index profiles which deviate substantially from

a step index form.

CONCLUSION

The work presented here provides a further interpretation of OTDR signatures for single mode

fibers. The technique provides a relatively accurate, non-destructive method to determine the

cutoff wavelength profile in most typical single mode fibers. Accuracies of ± 20 nm in this

prediction have been routinely achieved. Due to the relative speed and simplicity of the test,

the technique has a number of potential applications in such areas as manufacturing process

research and production process monitoring.

The theory behind this analytical approach assumes that the fiber tested exhibits ideal step

index behavior. Although a more detailed theory could be developed to include more realistic

fiber behavior, the step index model has shown good results. There are circumstances where

significant errors in the prediction may result. Therefore, care must be taken when attempting

to interpret the results of the technique.

REFERENCES

[1] P. Divita, V. Rossi; Backscattering Measurements in Optical Fibers: Separation

of Power Decay from Imperfection Contribution. Electronic Letters 15, p. 467,

(1979).

[2] M.S. O'Sullivan, R.S. Lowe; Interpretation of SM Fiber OTDR Signatures

Proceeding SPIE '86 Optical Testing and Metrology (1986).

[3] D. Marcuse; Loss Analysis of Single Mode Fiber Splices Bell System Technical

Journal, May-June (1977).

[4] L.B. Jeunhomme; Single Mode Fiber Optics: Principles and Applications Marcel

Dekker, Inc.); V. 4 p. 151 (1983).

125
40

POSITION (Urn)

Figure 1(a)

POSITION (km)

Figure 1(c)

m

a.Uio35

crz

(JM

in

300 2 3

POSITION (km)

Figure 1(b)

POSITION (km)

Figure 1(d)

8M

B

130

100 -

Figure 1(e)

POSITION (km)

Figure I Illustration of OTDR Analysis Techniqtio (Fiber A)

1(a) (b) Original OTDR signatures at 1308 nm and 1545 nm

1(c) (cl) Separated OTDR traces at 1308 nm and 1545 nm

1(e) Predicted cutoff wavelength profile correlated with

actual measurement performed using EIA-455-80.

126
Predicting The Cutoff Wavelength of Cabled Single

Mode Fibers From Cable Structure

W. H. Hatton and M. Nishimura

Sumitomo Electric Fiber Optic Corporation

Research Triangle Park, NC 27709

Introduction

Accurately predicting the cutoff wavelength of cabled single mode fiber

is desired by both fiber and cable manufacturers because fibers which exhibit

improved bend performance (longer uncabled cutoff) can be utilized. The

Electronic Industries Association (EI A) has adopted a test procedure (FOTP170) where cutoff is determined by measuring a 22 meter section of cable.

This method produces results which simulate worst case field conditions,

however, routine measurements would be uneconomical from a cablers point of

view, and impossible for a fiber supplier.

It has been suggested that a linear function which maps uncabled fiber

cutoff into cabled fiber cutoff be experimentally determined [1]. This is a

beneficial procedure, but without a complete understanding of how cable

structure effects cutoff, a different mapping would be required for each

cable design (a very time consuming process). In addition, it has been

proposed that cabling effects could be simulated by replacing the 22 meters

of cable by 22 meters of uncabled fiber looped at a minimum bend diameter of

280mm [2]. This technique may be adequate for certain cable designs, but in

most cases, the results would be extremely pessimistic due to an over

exaggeration of fiber curvature (i.e. effective cabled fiber bend diameters

of 100 mm to 200 mm are common). In this paper, a method for predicting

cutoff shift (uncabl ed-cabl ed) from cable design (effective bend diameter)

will be presented.

127
Experiment

To accurately determine cabled fiber cutoff both length and curvature

effects must be considered [3,4] . Changes in cable design primarily effect

curvature, and thus, the fiber deployment shown in Figure 1 was developed to

offer a variable scheme for simulating cable effects on cutoff. With the

addition of a variable mandrel, the fiber can be looped at a bend diameter

equivalent to that occurring inside the cable. The effective bend diameter

can be calculated from design parameters such as tube pitch, tube diameter,

and central tension member diameter (for loose tube designs). By adjusting

the mandrel size, the relationship between cable structure and cutoff could

be determined.

A collection of both matched clad and depressed clad single mode fibers

( 22 meter length) were prepared. Each fiber was subjected to various bend

diameters and cutoff was determined by comparing the spectral power

transmitted through the above deployment to that when an additional smaller

diameter bend was introduced. For each diameter, a linear least squares

routine was used to map uncabled cutoff into predicted cable cutoff. Figures

2A & 2B illustrates the relationship between cutoff shift (uncabled -

predicted) and effective bend diameter.

It is well known that effective cutoff of depressed clad fiber is less

sensitive to curvature effects [3], and to some extent this was reflected in

the measured results. There was, however, a critical diameter of 150 mm where

bending effects could not be neglected. Conversely, matched clad fiber, being

more bend sensitive [3], demonstrated a much stronger bend dependence. For

typical effective bend diameters of 100 mm to 200 mm, a cutoff shift of 40 nrn

existed. A model which neglects, or over exaggerate curvature effects, would

be inadequate for this fiber type.

128
Table 1 compares predicted cutoff shift to measured (FOTP-170) for

various loose tube cable designs and both fiber types. Figures 3A and 3B

shows fitted curves (predicted and measured) comparing uncabled and cabled

fiber cutoff.

Conclusions

Cabled fiber cutoff wavelength can be simulated accurately if both

length and curvature effects are considered. It has been shown that by

developing the relationship between effective bend radius and cutoff shift

(uncabled-s imulated cabled) cabled fiber cutoff can be accurately and easily

predicted. This eliminates the need for developing separate mapping functions

for each cable design. A more detailed comparison between predicted and

measured cabled fiber cutoff will be presented at the conference.

References

[1] J. P. Kilmer, IWCS Symposium Proceedings, pp. 49-56, 1987

[2] D. W. Peckham, Technical Digest, OFC 88, paper WE5, 1988.

[3] L. Wei, IEEE JLT Vol. LT-5, No. 9, p. 1147, 1987.

[4] W. T. Anderson, IEEE JLT, Vol. LT-2, No. 3, P. 238, 1984.

Table 1 Measured and predicted cabled fiber cutoff for various loose tube

cable structures

F i ber

Type

Effective Bend (mm)

D i ameter

Cutoff Wavelenqth (nm)

Sh i ft

Measured Pred i ct ed

Depressed 160 83 80

Depressed 146 89 87

Depressed 160 77 80

Matched 120 121 124

129
Light

Source 0

(A)

(B)

lm

Light

Source

76mm

i 20m Cable

i

i lm

76mm

2l

Variable Mandrel

lm

i

i

i 20m I

i 76mm

lm

Detector

Uncabled

Fiber

Detector

Figure 1 Schematic Diagram of A) deployment

for cable cutoff (FOTP-170), and B) simulated

cable cutoff.

Figure 2 Relationship between effective

bend diameter and cutoff shift (uncabledpredicted) for A) depressed clad fiber,

and B) matched clad fiber

Figure 3 Comparison between predicted and

measured cabled fiber cutoff for A)

depressed clad fiber, and B) matched clad

fiber

130
SIMULATED CABLE DEPLOYMENT FOR

CABLE CUTOFF WAVELENGTH MEASUREMENTS

J. P. Kilmer, R. M. Kanen, and W. T. Anderson

Bell Communications Research

Morristown, New Jersey

1. Abstract

The simulation of the cable cutoff wavelength deployment configuration with 22 meters of coiled

uncabled fiber is experimentally studied. The results show the simulation gives equivalent X ee results for

depressed-clad fibers and more conservative X,.,. results for matched-clad fibers.

2. Introduction

Cutoff wavelength refers to the wavelength below which a single-mode fiber is capable of supporting

more than one mode and is the primary limitation on the lower edge of the single-mode fiber's 1300 nm

operating window. The future use of multiple high bit-rate WDM systems around 1300 nm may make

the neglected 1250 to 1300 nm region more attractive than the 1550 nm region where there is greater

chromatic dispersion. Cutoff wavelength must be controlled in order to prevent single-mode system

operation in the fiber's multimode region [1-3]. The Electronic Industries Association (EIA) has two

definitions for cutoff wavelength. The first definition, fiber cutoff wavelength (X c/ ), is a convenient

parameter for a manufacturer to monitor for process control but provides the user with information

which is not directly applicable to system design. X^ is measured on a 2 m fiber sample deployed in a

fashion which is never found in the field, as described in the test procedure FOTP-80 [4] (see Figure 1).

The second definition, cable cutoff wavelength (X cc ), provides the user with a functional value of cutoff

wavelength which can be used in system design. FOTP-170 [5] describes the test procedure to measures

X ec

. CCITT Recommendation G.652 also contains similar definitions and procedures for cutoff

wavelength.

While the cabled fiber sample's deployment conditions can be specified in the detail specification,

FOTP-170 contains two possible default deployment configurations. The first default deployment

configuration is a 20 meter (65.6 ft.) cable length with one meter of exposed fiber at each cable end.

One 76 mm (3 in.) loop is inserted in each exposed fiber end (see Figure 2). 20 meters is considered a

worst-case minimum cable length likely to be deployed in the field (typically repair sections or stubs),

and splice organizers contain trays with space to coil slack fiber typically into 3 in. loops. Since this

deployment configuration requires 20 meters of cable, cable cutoff wavelength is by nature more difficult

to measure on a routine basis than fiber cutoff wavelength. Therefore, a variation on this cable

deployment configuration has been proposed in order to facilitate routine measurement. This second

default configuration consists of an uncabled fiber length totaling 22 m coiled in a loop of 28 cm

diameter or greater with the two 76 mm loops at the fiber ends (see Figure 3). The coiling of the fiber

length in a loop of 28 cm diameter or greater simulates the cabling effect. Cabling a fiber, if it affects

the fiber at all, may introduce some periodic bends which serve to reduce cutoff wavelength. Therefore,

uncabled fiber should give a slightly higher (or more conservative) \ ee value. Since the 3 in. loops are expected to dominate cutoff wavelength reduction [6], coiling the uncabled fiber in a loop of 28 cm or

larger should not affect the measured \ ee

.

3. Experiment

To verify that the 22 m of coiled uncabled fiber is an adequate simulation of the 20 meter cable length

with one meter of exposed fiber at each cable end, two cable samples were measured for \ cc 's according

to the first default configuration (Figure 2). One cable sample consisted of depressed clad fibers (DCF)

131
manufactured by the inside vapor deposition process (IVD) bundled together inside a crossply cable

sheath. The other cable sample consisted of matched clad fiber (MCF) manufactured by the outside

vapor deposition process (OVD) in a loose tube cable construction employing a reversing oscillating lay.

After measuring X cc in cable form, the fibers were carefully removed from the cable and deployed in the

coiled uncabled fiber deployment configuration (Figure 3). A comparison of the cabled and simulatedcable measurements appears in the following tables:

Coiled Fiber and Cable Cutoff Wavelength Comparisons (in nm)

DCF IVD Fiber in Fiber Bundle Cable

Fibpr

Blue/Blue 1173 1161 12

Til _ — /T~) Blue/Brown 1215 1212 3

Blue/olate lloO i i on lisy o-o Blue/ White iiy/ 1 OH7 120/ -10

Blue/rtea 1 1 QO 1 1 Q7 iiy / c -0

r>i av n Blue/Yellow 1180 1178 2

151.. „ /T*\ „ „L /~\ — ~ Blue/JJasn Urange ny4 1207 -16

Orange/Blue 1235 1218 17

Orange/Orange 1178 1168 10

Orange/Green 1177 1187 -10

Orange/Brown 1186 1199 -13

Orange/Slate 1176 1179 -3

Orange/White 1209 1212 -3

Orange/Red 1189 1189 0

Orange/Yellow 1187 1198 -11

Orange/Violet 1201 1215 -14

Orange/Dash Blue 1162 1171 -9

Orange/Dash Orange 1208 1196 12

Green/Blue 1204 1214 -10

Green/White 1213 1218 -5

Coiled Fiber and Cable Cutoff Wavelength Comparisons (in nm)

MCF OVD Fiber in Loose Tube Cable

Fiber Coiled Cabled Difference

Blue/Blue 1149 1147 2

Blue/Orange 1093 1099 -6

Blue/Green 1118 1094 24

Blue/Brown 1154 1117 37

Blue/Slate 1126 1112 14

Blue/White 1120 1109 11

Orange/Blue 1136 1128 8

Orange/Orange 1139 1116 23

Orange/Green 1129 1115 14

Orange/Brown 1150 1119 31

Orange/Slate 1118 1115 3

Orange/White 1120 1116 4

Green/Blue 1116 1124 -8

Green/Orange 1116 1098 18

Green/Green 1119 1088 31

Green/Brown 1130 1117 13

On average, there is a -2.65 ± 9 nm (1 standard deviation) bias for the IVD DCF fiber and a 13.68 ± 13

132
nm bias for the OVD MCF fiber between the two deployment configurations. While there is no

statistically significant difference between the two deployment configurations for the DCF fiber type,

there is a small bias for the MCF fiber type. This bias is expected due to the increased cutoff

wavelength bend sensitivity of the MCF fiber type [7]. Since the bias is positive in this case, the

simulated cable configuration gives a conservative (i.e., higher) value for X ec than the cabled

configuration.

4. Conclusion

The coiled uncabled fiber simulation gave equivalent X cc results to the cabled fiber deployment

configuration for a depressed-clad type fiber and gave slightly more conservative X ee results for a matched-clad type fiber due to the increased cutoff wavelength bend sensitivity of the MCF type fiber.

Consequently, the 22 m coiled uncabled fiber is an adequate simulation for cable cutoff wavelength

measurements. Since the coiled-uncabled fiber simulation is conservative, manufacturers can use it prior

to cabling to assure X cc requirements are met in the final cabled form.

REFERENCES

[1] F. T Stone, "Modal Noise in Single-Mode Fiber Communication Systems," SPIE Proceedings 500-

04, San Diego, (1984).

[2] R. W. Tarwater, M. J. Maslaney, D. L. Philen, F. T. Stone and D. G. Duff, "Bandwidth and

Cutoff Wavelength of Single-mode Lightguide System Considerations," IWC Symposium

Proceedings, pp. 261-265, Reno, Nevada, (1984).

[3] N. K. Cheung and P. Kaiser, "Cutoff Wavelength and Modal Noise in Single- Mode Fiber

Systems," NBS Special Publication 688, pp. 15-18, (1984).

[4] EIA-455-80, "Cutoff Wavelength of Uncabled Single-Mode Fiber by Transmitted Power,"

Electronic Industries Association, Engineering Dept., 2001 Eye St., N.W., Washington, D.C.

[5] EIA-455-170, "Cutoff Wavelength of Single-Mode Fiber Cable by Transmitted Power," Electronic

Industries Association, Engineering Dept., 2001 Eye St., N.W., Washington, D.C.

[6] J. P. Kilmer, W. T. Anderson, R. M. Kanen and C. M. Connor Davenport, "Cabling Dependence

of Optical Fiber Cutoff Wavelengths," IWC Symposium Proceedings, pp. 49-56, Arlington,

Virginia, (1987).

[7] V. Shah, "Effective Cutoff Wavelength for Single-Mode Fibers," Symposium on Optical Fiber

Measurements, NBS Special Publication 688, pp. 7-9, (1984).

133
R = 14cm

2m

Figure 1. Fiber Cutoff Wavelength (FOTP-80) Deployment Configuration

76 mm WWffl

Im

20m

Figure 2. Cable Cutoff Wavelength (FOTP-170) Default Deployment Configuration 1

,0- [) 76mm 76mm()

Figure 3. Cable Cutoff Wavelength (FOTP-170) Default Deployment Configuration 2

134
CUTOFF CHARACTERISTICS OF DISPERSION-SHIFTED AND

DISPERSION-UNSHIFTED FIBERS

V. Shah, L. Curtis, and W. C. Young

Bellcore

331 Newman Springs Road

Red Bank, NJ 07701-7020

(201)-758-3156

INTRODUCTION:

Dispersion-shifted (DS) single-mode fibers, having the zero dispersion wavelength in the low attenuation

1550 nm wavelength region, may be required to realize high capacity long haul transmission systems.

However, very little has been reported about the cutoff characteristics of such fibers. Recent work'1

! on

the cutoff characteristics of segmented core DS fibers reported the presence of a "hump" in the cutoff

region of the fiber. In this paper, we report the presence of such a "hump" in dispersion- unshifted (1300

nm optimized) single-mode fibers of both the matched as well as depressed-cladding design. Based on the

results of our study, a possible explanation of the presence of the "hump" in the cutoff region is presented

and resulting systems implications discussed

EXPERIMENTAL STUDY:

To measure the cutoff wavelength, the spectral powers, and PQ , transmitted through a 2-m long fiber

arranged in large and small single loops of diameters D = 28 cm and d = 5 cm are measured,

respectively. The normalized power (Pi/Pq) in dB is plotted as a function of the wavelength X, and the

cutoff wavelength is determined as the wavelength where the long wavelength edge of the curve increases

by 0.1 dB over the baseline, as per CCITT Recommendation G.652 and EI A FOTP-80. Fig. 1 shows the

normalized power-vs-wavelength curves for a segmented core DS fiber (curve A) as well as matchedcladding (curve B) and depressed-cladding (curve C) dispersion-unshiftcd fibers. As observed from Fig. 1,

the "hump" appears in the cutoff region of each of these fibers. To investigate the nature of the hump,

both Pj (dashed curve) and PQ (solid curve) are plotted as a function of the wavelength for the DS fiber

in Fig. 2. The "hump" of Fig. 1 appears in Pj(X) of Fig. 2. This indicates that contrary to the expected

135
decrease in the transmitted power due to rapid attenuation of the higher-order modes, an increase in the

transmitted power takes place in the wavelength region slightly below the cutoff wavelength. In the case

of DS fiber, one may associate this phenomenon with the cutoff of two higher-order mode groups ( LP&

and LPn ) having closely spaced propagation constants. However, in that case the cutoff curve is

expected to show two different slopes (both having same sign) in the cutoff region rather than the

"hump". In addition, the "hump" is observed in both depressed-clad and matched-clad dispersion

unshifted fibers which, unlike the DS fiber, do not have closely spaced propagation constants for higherorder modes.

A possible explanation of the "hump" can be offered in terms of coupling between the whispering gallery

mode'2

' formed in the cladding region due to the high-index coating surrounding the fiber and the

higher-order core mode. Power in the whispering gallery mode can couple to the higher-order core mode,

resulting in the "hump(s)". To investigate this possibility, removal of this coupling mechanism is required.

This can be achieved by removing the high index primary coating from the curved portion of the fiber

and very closely matching the cladding index by immersing the fiber in an appropriate index matching

oil. The whispering gallery mode is no longer supported under such conditions and therefore P\(\) should

show no "hump" in the cutoff region.

Fig. 3 shows the result of such an experiment carried out with the DS fiber. The figure shows Pi plotted

as a function of \. The solid curve corresponds to the fiber with coating and has two "humps" in the

cutoff region. The dashed curve corresponds to the fiber immersed in index matching oil and does not

have any "hump" in the cutoff region The results of these later measurements support the theory that the

"hump" in the cutoff region may be the result of mode coupling between the whispering gallery mode

and the higher order core mode.

CONCLUSION :

In conclusion, the fibers investigated (both DS and dispersion unshifted) exhibit a "hump" in the cutoff

region. Experimental results are also presented which support the theory that these "hump(s)" occur as a

result of coupling between the whispering gallery mode and the higher order modes.

136
REFERENCES:

[1] V. Shah, G. Kinshofer, L. Curtis, and W. C. Young, Tech. Dig. OFC88, New Orleans, Paper WI5,

(1988)

[2] A. J. Harris, and P. F. Castle, J.of Lightwave Tech., Vol. LT-4, No. 1, pp. 34, (1986)

CD

o

or

UJ

o

Ql

Q

LUN

DC

O

1000 1200

WAVELENGTH (nm)

1400

Fig. 1 Cutoff curves for a segmented core DS fiber (curve A), a matched-cladding fiber (curve B), and a

depressed-cladding fiber (curve C).

137
DC

LU

§ en

Q

LU

^

>

1=2

I*is

GC

5 ci

I |_ \

\

28 cm

4cm

1 1

1000 1200 1400

WAVELENGTH (nm)

Fig. 2 Pi(k) (dashed curve) and PQ (k) (solid curve) plotted as a function of X for a DS fiber.

(0

c 5

>»

(0

n

<

o

5

o

Q.

"O

<D

*-> E

(/) c

CO

\

3 -

1000

1 1 1

With Primary Jacket

With Index Matching Oil

1 1

1100 1200

Wavelength (nm)

1300 1400

Fig. 3 Pi(X.) plotted as a function of X for a DS fiber.

138
1300 nm fiber characterization by LP0i and LP02 interferogram

equalization wavelength measurement

Francois Gonthier, Suzanne Lacroix, Richard J. Black and Jacques Bures

Laboratoire d'optique guidee et d'opto-electronique, Ecole Polytechnique, Departement de genie physique

C. P. 6079, Succ. A, Montreal, Quebec, H3C 3A7, Canada

Abstract: When operated in the visible wavelength range with an on-axis launch, a standard telecommunications

fiber supports the first two circularly symmetric modes LP01 and L?o2- The wavelength response of such a modal

interferometer is observed and permits the measurement of the group velocity equalization wavelength, which is used

to characterize the fiber parameters in the case of near-step-index. For more complex profiles, this measurement can

provide valuable additional information regarding the form of the profile.

Introduction

Telecommunications fibers, which are single-mode at 1300 nm, are few-mode in the visible and very-near

infrared. The interference of those modes can be used to characterize the basic parameters of the fiber. For instance,

McMillan et al.[l] proposed to use the observation of the equalization wavelength the LP01 and LPn modes to

determine precisely the V-parameter of the guide as a way of replacing the often biased measurement of the second

mode LP11 cutoff wavelength Xq. Since it is a measurement of the phase between the modes rather than an intensity

one, the determination is less sensitive to the bending and length of the fiber. However, due to the symmetry of the

modes involved and the wavelength at which the LPoi-LPn equalization occurs, this experiment is delicate to

implement.

In this paper, we propose to use the LP01-LP02 group velocity equalization wavelength in the

characterization of telecommunications fibers in the visible domain, where they can be used in sensors applications

[2,3]. The observations are made using an all-fiber modal interferometer. The step-index profile is examined and more

complex profiles are discussed.

LP01-LP02 modal interferometers

When operated in the visible wavelength range, single-mode telecommunications fibers (typically V « 2.2 at

1300 nm) become few-mode fibers (V - 4.4 at 650 nm). To realize a LP01-LP02 modal interferometer - i.e. a fiber

in which only LP01 and LP02 propagate and beat - one can splice strands of single-mode fiber (V « 2.2 at 650 nm) at

each end of the telecommunications fiber under test. If the two splices are well centered, only the modes of circular

symmetry LP01 and LP02 are excited and their interference is detected with an optimum contrast.

This structure is well described by a "coupling-beating-coupling" model [2,3]. The coupling is ensured by the

splices where there is an overlap between the upstream guide modal fields and the downstream ones. In between, the

excited LPqi and LP02 modes propagate without coupling and accumulate a phase difference <|) due to their different

139
propagation constant 0oi anti Po2- Given L the few-mode fiber length, we have <|> = 60 L, where 60 = 0oi - P02 k

a function of wavelength.

The total transmission T of such a modal interferometer takes the form:

2 2

T = P x +P 2 + 2 P x P 2cos <|»

where Pi and P2 are the fractions of initial power transmitted in the few-mode fiber through LPrji and LP02

respectively. Note that Pi and P2 vary slowly with wavelength compared to cos <|> and depend on the fibers

parameters.

Equalization wavelength: Theory and measurements

The equalization wavelength is, by definition, the wavelength at which the group velocities of two given

modes, say LPrji and LPo2, are equal. Thus, at this particular wavelength, there is no intermodal dispersion. It can

be shown that it corresponds to a maximum of the propagation constants difference 60 and thus of the phase <|>.

Consequently, the equalization phenomenon appears in the modal interferometer ^.-transmission as an achromatic

fringe, i.e. a fringe which shows a stationary dependence on the wavelength (d<j)/dX = 0).

The experiment uses a white light source to input power in the fiber. The output end of the fiber illuminates a

monochromator equiped with a photomultiplier. To normalize the signal, the system response is first recorded using

only the 650 nm-single-mode fiber. Then, the fiber is cut into two halves and the few-mode fiber under study is

spliced in between. The modal interferometer response is then recorded and normalized.

Figs. 1 and 2 show typical normalized responses for two different index profiles and fiber lengths L «» lm. In

each case, the achromatic fringe position can be determined with a very good accuracy (± 0.5nm). As opposed to

fiber cutoff wavelength measurement and as in the case of the LPoi and LPn modes [4], the LP01-LP02

equalization wavelength position does not depend on the fiber bends (Fig. 2 a) and b)) or length. Reproducibility is

about ±1 nm.

Index profile dependence of the LP01-LP02 equalization wavelength

The LP01-LP02 equalization wavelength, as does the cutoff wavelength, depends on the fiber index profile. In

the more simple case (matched-cladding step-index), the equalization takes place for a constant value of Ve = 4.44.

The equalization wavelength \q can thus be easily related to the LPn cutoff wavelength Xc, knowing that the

normalized frequency at cutoff is Vc = 2.40.

kc = (Ve / Vc) = 1.85

Furthermore, one can fit the response of a theoretical perfect step-index fiber on the experimental graph, as

was done in Fig. 1. This gives an equivalent profile that has a radius of 4.92 ± 0.01 pirn and a index step of 0.00312

± 0.00001. The errors are due to the precision in the measurement of the interferometer length.

For graded-index profiles (i.e. clad power-law profile characterized by the q-parameter), Ve and Vc augment as

the profile tends to a triangular one (q = 1). For a near-step-index profile (large q), the ratio Vc /Ve stays close to its

140
640 650 660 670

Wavelength (nm)

680 690

Fig. 1: Experimental wavelength response of an LP01-LP02 modal interferometer made with aim long

step-index matched-cladding fiber and its theoretical fit with an equivalent step-index (radius = 4.92 [im,

index step = 0.00312). The equalization can be seen to occur at 666 nm.

0.5

c

o

*5S

B 04

e

0.3

0.42

e

1 0.40 h

S

ine

g 0.38

0.36

(a) unbent, depressed-cladding

(b) bent, depressed-cladding

670 680 690 700

Wavelength (nm)

710 720

Fig.2: Experimental wavelength responses of an LP01-LP02 modal interferometer made with aim long

depressed-cladding fiber . Although the fringe contrast decreases in going from (a) unbent fiber to (b)

fiber bent with a radius of 1.7 cm, the equalization wavelength remains close to 699 nm.

141
1.85 asymptotic value. Fig. 3 shows how it varies for q>10, together with the theoretical LPn wavelength cutoff

for the Fig. 1 fiber. As a result, the cutoff wavelength = 1233 nm, deduced from the Xq measurement (666 nm),

assuming a perfect step-index, may be underevaluated (by 7 nm if q = 20, and 2 nm if q = 40).

More complex profiles have different effects on the equalization wavelength. While grading of the profile

decreases Xq away from Xq, the effect of a cladding depression has the opposite. As an example, assuming perfect

index steps, the Fig. 2 depressed-cladding fiber theoretical LP01-LP02 equalization occurs at Ve = 4.8 and the LPn

cutoff at Vc = 3.0 (Ve / Vc = 1.6).

10 20 30 40 50 60 70

Power-law profile parameter, q

E

f ~

c E

1

5

9 << 3 w O

Fig. 3: Ratio of V values for LP01-

LP02 equalization and LPn cutoff

for the standard near-step-index

power law profiles. Note that, for

q>3, Vg/Vc - 1.848 + 0.168 / q +

1.796 / q

2 - The right hand axis

gives the LPn cutoff wavelength

deduced from the equalization

wavelength measured for the near

step-index fiber of Fig. 1.

Conclusion

The measurement of the equalization wavelength of the group velocities LP01 and LPo2> is simple, rapid and

very precise. For a fiber which has a matched-cladding step-index profile, it permits an accurate measurement of the

LP11 cutoff wavelength which is independent of fiber bends and length. For near step-index fibers, the error on the

cutoff wavelength evaluation is small and is systematically negative. Furthermore, a theoretical fit of the wavelength

response permits the determination of a precise step-index equivalent for the visible wavelength domain. For more

complex profiles, combination of the equalization measurement with other standard measurements provides

information about the form of the profile.

[1] J.L. McMillan and R.C. Robertson, "Cutoff wavelength determination in single-mode optical fibres by

measurement of equalization wavelength", Electron. Lett. 20, pp.698-699 (1984).

[2] F. Gonthier, S. Lacroix, F. Ladouceur, R.J. Black et J. Bures, "Interferometres modaux a symdtrie circulaire

dans les fibres optiques", submitted to Annales des telecommunications.

[3] S. Lacroix, F. Gonthier, R. J. Black, and J. Bures, "Tapered-fiber interferometric wavelength response: the

achromatic fringe", Optics Letters 13, pp.395-397 (1988).

[4] A.C. Boucouvalas, S.C. Robertson, and J. Walker, "Optical fibre far field ESI determination using the

equalisation wavelength", Technical digest, Symposium on Optical Fiber Measurements pp. 137-140 (1986).

142
FIBER GEOMETRY. A FIBER MANUFACTURER'S VTEWPOTNT

by David A. Hall and Robert W. Smith

Corning Glass Works

Introduction:

The measurement of optical fiber in the manufacturing environment generally falls into one of three

catagories: fiber loss measurements (attenuation and point defects), information capacity (bandwidth and dispersion) and fiber geometry (the measurement of physical dimensions of the fiber or

its refractive index profile). Much has been written and discussed about loss and information

capacity measurements, but much less is heard about the measurement of fiber geometry. In this

paper the considerations for evaluating physical fiber geometry measurements and the experience

we have had at Corning Glass Works in implementing fiber geometry measurements in manufacturing will be discussed. It will be shown that available fiber geometry measurement capability

meets current manufacturing requirements, but that potential changes in fiber specifications will

require substantial improvements in measurement technology.

Considerations for fiber geometry measurement:

There are many parameters which must be considered when selecting a fiber geometry measurement approach or implementing a system in the manufacturing environment. These considerations

fall generally into three catagories: measurement performance, system industrialization, and commercial and support issues.

The performance of a measurement system is the most obvious issue in determining suitability, but

it may be one of the most difficult issues to truly resolve. The parameters we consider as performance parameters are the system's accuracy, reproducibility, speed and conformance with standards.

We consider accuracy to mean the degree to which the measurement represents true calibrated metrology, referenced to some acknowledged external standard. Reproducibility is similar to repeatability; but whereas repeatability is the ability to get the same measurement result with repeated

measurements such as a short term cycle test, we use reproducibility to indicate the long-term

stability of the measurement system as well as repeatability. That is, reproducibility is the ability to

get the same measurement result with a complete new measurement set-up days or weeks later.

Speed is a measure of the system's measurement time, both the actual cycle time (measurement

143
sequence only), and the likely total measurement time including operator set-up. Standards conformance is the degree to which the measurement algorithms and set-up meet the requirements of the

national and international measurement standards recommendations.

The performance parameters must be evaluated against the true requirements of the system, since

some parameters will probably trade off against others (measurement speed and accuracy, for

instance). Also there may be a substantial difference between specified performance and demonstrated performance under factory use conditions. The way in which performance parameters are

specified and the way tests are conducted and evaluated can have a significant bearing on the usefulness of the data. Demonstrated performance also should include consideration of the skill and

training of the operator, particularly when targeting a system for use on the factory floor.

Industrialization parameters are fairly general in nature and therefore much more subjective in

evaluation. These include the nature of the operator interface, reliability and maintainability,

flexibility and integration.

In a large manufacturing facility the operator interface may be more important than in a small

operation where one person may be operating the system. The operator interface includes the ease

of use and accessibility of the hardware, the software interface and degree of automation, the

sensitivity of the measurement performance to operator set-up and operation, and the level of training required. Operator actions and judgement should have very little influence over the accuracy

of the measurement results in a factory environment.

Reliability and maintainability are difficult parameters to evaluate unless information is available

from other users of a system. One can subjectively evaluate the system for anticipated maintenance

problems, complexity and the use of proprietary parts and assemblies. In an automated system,

the use of proprietary CPUs and system architectures is a concern. The availability of testing and

calibration devices and maintenance training is also important

Flexibility and the ability to integrate the measurement into the manufacturing system are both

important considerations. Flexibility allows one to easily change the measurement parameters or

sequence, and even the measurement algorithms, if a measurement standard or product parameter

should change. In a highly integrated manufacturing environment, the ability to automatically trans

fer data to manufacturing systems and deal with ware flow management systems is also important

144
The commercial issues deal with the supplier relationship, and include value, warranty, support

and training, spares and maintenance agreements and upgrade policies. Having a good working

relationship with the system supplier is very important, and can impact the total operating cost of

the system.

Coming's Experience with Fiber Geometry Systems:

The following information describes Coming's experience in evaluating and implementing fiber

geometry systems for factory floor operation. The opinions and decisions discussed are in the context of Coming's requirements and are not meant to speak to other requirements or operations.

The background of our system choice, the general structure and operation of our system, and the

measurement results we currently get with the system will be discussed.

When the project to develop a system for factory measurement of fiber geometry began, the only

commercial system available was an image shearing microscope system. One of these was purchased and is still in use today, but it was not adaptable to our manufacturing operation. While it was found to be quite accurate and very flexible, it was very difficult to use well and was quite

slow. At that time it had no automation capability, so it would not integrate with our manufacturing systems. Other imaging systems which have been evaluated are much closer to meeting our

requirements, but most are based on proprietary architectures and hence are very inflexible. The

refractive nearfield measurement system used in the plant for process analysis has proved to be

very difficult to use as a production measurement.

With this background, it was decided to develop our own geometry measurement system in conjunction with an electro-optic measurement system company. The system is based on gray-scale

image analysis technology, using a video camera and scan digitizer hardware, custom optics and

illumination system, and a customized PC based computer system. The software provides

machine sequence, data acquisition and analysis, operator interface and communication of measurement results to the manufacturing system. This measurement system provides a fast, fully automated measurement which meets the performance requirements. It is flexible at both the measurement parameter and programming level, and is fully integrated with our manufacturing system. An

additional benefit of a custom system is that the operator interface can be made similar to other

measurement systems, reducing training requirements significantly. The principal disadvantage of

a system such as this is the cost of development.

145
The implementation of this system in our manufacturing operation, and the results which we obtain

with the system will now be discussed. Reproducibility will be separated from accuracy, as we

consider reproducibility to be controlled by the system and its operation on the factory floor, while

accuracy is primarily a function of the quality architecture in place.

The operation of the fiber geometry systems begins with fiber end preparation. A commercial fiber

cleaver and an interference type end inspection system are used. All fibers are required to be

inspected, and to have < 2° end angle and a clean cleaved surface. Since the end quality directly

affects the measurement results, careful management of end preparation is very important.

The operator interface of the measurement system is designed to help the operator mount and align

the fiber and enter the required information. The system is automated so that it completes the measurement from'final auto-focus through results reporting without further operator intervention. The

measurement algorithms are designed to conform with CCl l'l and ELA requirements for data fits

and computations. The system provides clad diameter, clad non-circularity and core-clad concentricity measurements on single-mode fiber; and all of the above plus core non-circularity on multimode fiber. The measurement time for this system is approximately 1 minute machine cycle time.

A significant development unique to this instrument is an illumination system designed to enhance

the contrast of the image and to improve the consistency of data acquisition. The fiber end is illuminated in such a way that the fiber cladding is dark against the bright background and bright core,

producing a gray-scale image with very steep edges. This technique sharpens the contrast, thereby

improving the dimensioning precision on the edges. It also reduces the effect of the chips and

hackle on the fiber surface, since the image is like a shadow of the fiber end.

The following results for clad diameter measurements show excellent system reproducibility.

Ongoing certification results using multiple fibers in a production environment show system reproducibility to be 0.15 }i.m (1 sigma). This measure of variability includes operator interface and setup, fiber end preparation, and system variability over many days. The measurement system itself

in a short term cycle test on a single fiber end exhibits a repeatability of 0.08 - 0. 1 (im (1 sigma).

The actual resolution of the detection system on an image edge is 0.05 (im.

The quality systems required to ensure accuracy were much more complicated and difficult to

develop, since they involved requirements for traceability and standardization. These considerations involve definitions of sub-micron level measurements and edge and illumination effects.

NBS studies have shown that the ability to relate different optical measurement techniques is

146
limited to around 1 (J.m, even on chrome-on-glass standards. Since our production measurement

system is designed specifically to measure optical fiber, a fiber was chosen as the most useful

calibration standard.

The image shearing microscope at our product engineering laboratory is used as our standardization system. This system was carefully calibrated using a recticle traceable to National Physical

Laboratory, and several selected fibers were measured with the system many times. The same

fibers were then measured with both a photographic recticle overlay technique and a full contact

measurement performed by Eli Whitney Metrology Laboratory. Comparison of all three measurement techniques showed agreement to within 0.2 |im, indicating accurate calibration of the reference system to that level.

Calibration begins by selecting fibers for their dimensional consistency and targeting them by measuring them several times on the calibrated reference system. The production system is then calibrated by measuring the targeted fibers and setting a calibration parameter for agreement with the

targets. A certification system is also used to ensure proper continued operation of our measurement systems. This system works by requiring the measurement of targeted fibers daily by operators and an ongoing computerized statistical analysis of the resulting data for bias and reproducibility. If a measurement system fails to conform to the certification requirements, it is shut down and

troubleshooting is done until the problem is resolved. Using this calibration and certification system, our factory floor measurements are accurate to within 0.5 |i.m. This is based on a ± 0.2 \im

reference accuracy and a 95% confidence interval (2 sigma) of ± 0.3 |im. This capability is adequate to support a product specification of ± 2.0 |0.m clad diameter tolerance, using product assurance statistics.

Future Directions and Requirements:

The directions we see currently for factory fiber geometry measurements are both good news and

bad news. The good news is that measurement equipment manufacturers are developing and

announcing new and upgraded fiber geometry measurement systems, some of which are aimed at

the factory floor. Since it is usually more cost effective to buy than build equipment if it meets the

requirements, this is a very positive sign. The bad news is that fiber geometry specifications and

therefore requirements for measurement systems are getting tighter. Recent market requirements

have driven diameter tolerances down from ± 3 Jim to ± 2 fim and there is interest in even tighter

specifications. Measurement systems and calibration capabilities will have to be substantially better

if fiber specifications continue to tighten.

147
Several future needs for fiber geometry measurement systems follow from the forgoing information. First we need continued improvement in fiber end preparation and end inspection systems.

The current end inspection systems work, but are tedious to use and difficult to interpret. A

system which would evaluate an end automatically and would maintain statistics on a cleaving

system would be very useful in insuring the quality of end preparation for fiber geometry and other

measurement systems. A reduction in end preparation time is also quite important in the manufacturing environment

Second, we need to continuously improve the capability of our measurement systems themselves.

As was indicated above, the current measurement capability of available systems is not adequate

for assuring conformance to sigficantly tighter fiber geometry specifications. These improvements

need to be not only in measurement precision, but in automation and stability of the systems as

well.

Finally, the standards and metrology community should develop and provide calibration and standardization systems as the basis for ensuring agreement within the optical fiber community on fiber

geometry measurements. Standardized measurement procedures and techniques are only part of

the solution. Transfer standards and calibration methodologies are also needed. As we move into

the subscriber loop era where inexpensive mechanical splices and connectors are required, calibrated sub-micron measurements will be imperative in order to ensure the compatability of various

components.

Summary:

The development of a fiber geometry measurement system and the required quality systems to

support it at Corning has resulted in an excellent production measurement capability which meets

the requirements of current fiber specifications. This was done as an internal development project

to meet internal requirements for process control and product assurance. As more vendors of

measurement equipment and more users of the measurements emerge, it is evident that the optical

fiber business needs a mechanism for standardization and calibration of sub-micron level fiber

measurements. As the business requirements change and require tightening of product specifications, this need will become even more acute.

148
Non-Contact, Interferometric Determination of the Outside

Diameter of Optical Fibers

M. J. Saunders

AT&T Bell Laboratories

2000 Northeast Expressway

Norcross, Georgia 30071

To measure the outside diameters of optical fibers, AT&T Technologies uses a forward

scattering method111 pl requiring a calibration fiber of known outside diameter. The diameter of this

fiber was previously measured with an interferometer to determine the distance that two originally

contacting plates are separated to just make contact with the fiber placed between the plates. [3]

Therefore, it is subject to errors caused by the presence of small foreign particles, and to the uncertainty

in determining when contact occurs. A new, accurate, (0.1-0.2 Jim) non-contacting measuring technique

to determine the outside diameter of the calibration fiber is the subject of this paper.

The fiber to be measured, A, is placed over a hole in plate B (Figure 1) that is supported by

an uncoated optical flat, C. The interferometer, D, is a commercially available Michelson interferometer

attached to the body tube, E, of a microscope that has a fine focus least count of 1 p.m. This

microscope is attached to a cross-slide that serves as a coarse focus control. The beamsplitter cube, F,

together with other optical elements, form a 10X microscope objective. Therefore, the focussed object is

crossed with interference fringes.

A white light source, G, is used to illuminate the interferometer that is adjusted until the

optical path Y0 between the center of the beam splitter, F, and the uppermost point on the fiber is equal

to L0 , the invariable path between the center of F and H, the reference mirror of the interferometer

(equal distance condition). Under this condition, a straight, uncolored interference fringe appears on the

uppermost part of the fiber and, by adjusting the tilt of H, this fringe is oriented parallel to the fiber axis.

(That only one fringe appears is due to the curvature of the fiber.) As the interferometer is moved to

upset the equal distance condition, the uncolored fringe is replaced by one of color, whose brightness

decreases with the deviation from the equal distance condition, and if this condition is upset by only 0.3

\un, the brightness decrease is discernible to the eye.

149
Bracket, J, attached to the microscope, supports a retroreflector, K, that is the movable

mirror of another interferometer contained within the dashed circle on Figure 1 where M and L are

mirrors, N is a laser and detector and O is a bidirectional fringe counter used to determine the motion of

K and D. The least count of this interferometer is 0.16 |xm.

After the equal distance condition has been obtained with the fringes on the top surface, the

counter is zeroed and D lowered until another set of fringes, formed on the bottom surface of the fiber,

is seen and the interferometer is adjusted so that the equal distance condition again holds. (Yi = the

new distance between the center of the beamsplitter and the uppermost part of the fiber.) From Figure

2, equating the optical paths from the beamsplitter to the reference mirror and to the lower surface of the

fiber,

2 |yx + nCL (D 0 -D) + n-(D- Dco) + n

+ Dco| = 2 Y0 , (1)

where /icl. n

~ and n

+ and D0t D and Dco are the group indices and the diameters of the outer cladding

(fused silica), depressed cladding and the core, respectively.

Solving for D0 ,

Yn — Y\ n — tier i

+ — ricr D0 = — - DA" + Dco (A" - A

+

) , where A~ = — and A

+ = — . (2)

Using values characteristic of a depressed cladding design for single mode fibers,

D = 55 \im, Dco = 8.2 \im, A" = -1.6 x 10"3 and A

+ = 3.3 x 10~3

, the terms in eq. (2) other than

Y0 - Yx are evaluated as 0.05 urn. Consequently, calculating DQ from the following equation (2^ will

give fiber diameters too small by the systematic error of 0.05 \un,

Z>„ = ^i-. (2)

As the interferometer is moved from the position for fringes on the uppermost surface of the

fiber to that for fringes on the lowermost surface, monochromatic interference fringes counted by O, give

(Yo-Yi) of equation (20.

150
The group index, nCL, is determined by the method of Figure 3. The fiber is replaced by

two plane parallel, optically contacted, fused silica discs (A, B) of thickness TA and TB , respectively.

The central white light fringe is obtained alternately along paths 2 and 3 by lowering and raising the

interferometer. Y2 and Y3 are the distances between the center of the beamsplitter and the top of A and

B when fringes are formed off the bottom of B along paths 2 and 3 respectively. The reading of the

counter is equal to A.

The thickness, TA> of disc A is equal to the interferometer motion required to obtain the

equal distance condition for fringes formed on the top surface of disc A after having first obtained this

condition for fringes on the top surface of disc B. Then, using the L0 defined in Figure 2, and equating

Lq with the optical paths along paths 2 and 3, one obtains

A + TA nCL = rp . (3)

Fifteen measurements of TA and A give nCL = (1.4810 ± 0.0003), in agreement with the group index

calculated for fused silica at X = 0.58 jjm. [4] Ten measurements of the outside diameters of each of 12

fibers were determined (total error -0.1 |im) and the average values compared with the forward

scattering results. The average difference between the two methods is 0.2 Jim, the largest positive and

negative differences being + 0.6 \xm and - 0.4 ^im, respectively. These largest differences could be due

to rotating the fiber about the fiber axis in one method (forward scattering) and keeping it constant in the

other.

I thank W. B. Gardner, R. B. Kummer and D. W. Peckham for very helpful discussions, and

C. Bice for making the forward scattering measurements.

151
REFERENCES

1. L. S. Watkins, /. Opt. Soc. Am. (64), 767 (1974).

2. D. H. Smithgall, L. S. Watkins and R. E. Frazee, Jr., Applied Optics (16), 2395 (1977).

3. D. H. Smithgall and C. M. Schroeder, Symposium on Optical Fiber Measurements, 1980, Boulder,

Colorado, October 28-29, 1980.

4. F. L. Kapron, Electronics Lett. (13), 96 (1977).

FIGURE 1

SCHEMATIC DIAGRAM OF APPARATUS

FIGURE 2

DIAGRAM TO DETERMINE OPTICAL

PATH THROUGH FIBER

y2 ®nciV3

®nci -LB

FIGURE 3

SCHEME FOR DETERMINING

GROUP MDEX n d

152
ON THE DEVELOPMENT OF A CALIBRATION STANDARD

FOR FIBRE 6E0METRY MEASUREMENT

1 2 1 J.G.N. Baines, A.G.Hal lam, K.W.Raine

1 - National Physical Laboratory, Teddington, Middlesex, UK

2 - York Technology, School Lane, Chandlers Ford, Hants, UK

Introduction.

The optical -fibre industry is now sufficiently advanced as to require traceable

and transferable calibration standards -for both optical -fibre manufacturers and

system designers. Improvements in optical -fibre -fabrication techniques and the

need -for tighter tolerances in splice technology have generated a requirement

-for high accuracy measurement of -fibre geometry.

This paper reports on investigations into the development o-f a chrome-on-glass

mask suitable -for geometrical calibration o-f re-fr acted near-field (RNF)

instruments. The mask takes the -form o-f a clear circular aperture, 125um in

diameter, surrounded by opaque chromium. Chrome-on-glass artefacts have been

investigated because of their convenience of use, stability and longevity. In

addition, dimensional measurements of such objects are well understood since

they are frequently used in the microelectronics industry for calibration

purposes, (1).

The use of RNF for Geometry Measurements.

Refracted near-field (RNF) instruments determine fibre geometry from the

coordinates of points corresponding to 50% of the refractive index step at the

oil/cladding or cladding/core boundary. In the case of graded-index cores some

arbitrary level (eg 5%) is used. A computer program is used to calculate the

geometrical properties such as average diameter, non-circularity and

core concentricity error.

The normal mode of operation of the RNF instrument is analogous to the idea of a

scanning spot phase-contrast microscope. However the transmission object

considered in this work is unlike an optical fibre, which is a refractive

object, and may introduce different systematic effects into the measurement.

These effects have therefore been investigated.

Measurements.

The chrome aperture (nominally 125um in diameter) and the cladding diameter of a

number of optical fibres were measured using an interf erometrical ly calibrated

image-shearing microscope with a numerical aperture (NA) of 0.45. The overall

uncertainty was +/-0.3um. When these fibres were used to calibrate an RNF

instrument (launch NA 0.85, blocking NA 0.47) it was consistently found that the

chrome aperture was measured by the profiler to be about 0.5um smaller than when

measured by image-shearing. Similarly, if the aperture was used to calibrate the

RNF instrument this led to fibre diameters being measured about 0.5um too large.

However a similar mask used on a second RNF instrument having lower launch and

collecting NA's (0.54 and 0.41 respectively) did not exhibit this effect,

153
although, as can be seen in the -following section, this result was incidental

and due to the particular operating conditions used.

Diffraction Effects.

The measurements show that the use o-f a chrome-on-glass mask -for the geometrical

calibration o-f RNF instruments can lead to a systematic error. This arises

because as the opaque chromium -film cuts across the -focussed spot, the reduction

in size of the unobscured light spot -forces it to di-f-fract through larger

angles. Since the collecting NA is greater than the launch NA the amount of

light collected when the -focussed spot is initially intercepted will be

proportionately larger than expected. This is because some o-f the light which

would have been intercepted by the blocking disc is now diffracted past it. The

magnitude of this effect will clearly depend on the difference between the

launch and collecting NA's. As more of the spot becomes obscured the diffraction

angle soon reaches 90 degrees and the collecting optics is no longer able to

capture all of the light. The resultant reduction in the measured light power

from that expected from geometrical arguments is dependent on the collecting NA

and possible changes in the angular light distribution which could provide some

compensation. Thus there are two opposing mechanisms which can distort the edge

profile. The overall effect is not easy to calculate but is obviously determined

by the launch NA, collecting NA and blocking NA. For example, if loss of light

from the collecting optics is a dominant factor then an opaque chromium disc

will appear too large whereas a clear aperture will be measured too small.

To demonstrate this, clear and opaque discs nominally 125um and 60um in diameter

respectively were measured as a function of blocking NA. The results are shown

in Fig 1. It can be seen that the smaller blocking NA leads to smaller diameters

of the aperture (dashed curve), while the opaque disc shows the opposite effect

(solid curve) . This is because in this RNF instrument the collecting NA is only

slightly larger than the launch NA. Thus, when the blocking NA is small the

amount of light gained through diffraction into the full collecting NA is a

proportionately smaller fraction of the light collected than when there is no

obscuration. The loss of light from the collecting optics therefore appears the

dominant effect. Clearly the blocking NA can be increased to a point where this

is no longer the case as can also be seen from the Figure when the blocking NA

exceeds about 0.4.

Focus Effects.

A 125um clear aperture was mounted on a hollow hemispherical profiling cell and

inserted in the first RNF instrument. The numerical apertures for the launch

objective, blocking disc and collecting optics were 0.85, 0.47 and 0.83

respectively. A raster scan was performed over the region of the aperture with a

scan increment of 2um, and a least-squares fit to the 50% intensity-level data

points was applied to obtain the average diameter.

Fig 2 shows the effect on measured diameter of varying the focus position. The

diameter scale was calibrated against a fibre measured on the microscope,

discussed above. The optimum focus position was taken as the edge over which the

approx 20%-80% risetime was minimum. A positive focus offset indicates the

objective being withdrawn from the fibre end. The approximate sensitivity to

focus is +0. 12um diam/um focus.

A series of RNF scans across the edge of the aperture is shown in Fig 3, where

increasing optical signal is in the downwards direction. A slight asymmetry with

focus can be seen along with a discontinuity in the slope of the edge. Spherical

154
aberration effects appear to be evident since the in—focus position should

correspond to a minimum or maximum i-f di-f -fraction were the only difficulty.

The pro-filing cell was then -filled with matching oil o-f index 1.47. This had the

e-f-fect o-f reducing the exit NA from the cell, which thus increased the

proportion o-f light that was blocked and reduced the sensitivity to di-f-f raction

losses at the collection NA. The measured diameter was -found to be slightly

larger -for the wet cell, as expected -from the preceeding argument; by about

0. 2um. The -focus sensitivity was +0.29um diam/um -focus. Typical edge plots are

shown in Fig 4. The marked 'knee' in the slope o-f the edge that occurs -for

positive -focus offsets is not fully undestood but its asymmetry with focus

position is probably due to the interaction of diffraction effects with

spherical aberration in the objective. However the discontinuities in the slope

were found to disappear when a large area detector was positioned immediately

behind the mask in place of the normal collecting optics.

It was found that by fitting a 0.9NA objective with a coverslip correction ring

the 'knee' could be made to occur for negative focus offsets by adjustment of

the correction ring. The slope of the focus sensitivity was also reversed so

that a positive focus offset gave a decrease in measured diameter. For a

particular correction ring setting the 'knee' occured at the in-focus position,

in which case the diameter showed reduced sensitivity to focus position.

However, the in-focus diameter was found to be relatively insensitive to the

level of coverslip correction.

Discussion.

The problems associated with the use of an aperture in chrome for calibration of

the RNF profiler have been discussed. Due to diffraction effects the measured

diameter of the aperture is dependent on the particular configuration of launch

NA, collection NA and blocking NA, as well as focus position; in the latter case

spherical aberration effects the result also. The use of an aperture for

calibration is therefore only valid under strictly controlled conditions.

A chrome aperture of annular form is under development which is expected to show

less sensitivity to diffraction effects. An average of the internal and external

diameters of the annulus will constitute the reference dimension since this will

have reduced sensitivity to axial ly symmetric aberrations, (2) , blocking NA and

focus. This is because line-spacing rather than line-width will be the

measurement parameter. However, as in high resolution microscopy, the presence

of an asymmetric aberration such as coma will be detrimental.

Finally, similar calibration discrepancies have been observed in the scanning

beam method for fibre diameter measurement. This technique is often employed for

control purposes during fibre pulling. Calibration is usually by means of a

series of opaque wires of known diameter. Edge response effects in this

environment are under investigation.

References.

(1) M.J. Downs, N.P.Turner, "Application of optical microscopy to dimensional

measurements in microelectronics 11 , SPIE vol 368, 1983 (Microscopy - Techniques and Capabilities).

(2) CP. Kirk, "Aberration effects in an optical measuring microscope", Appl Opt.

vol 26, no 16, pp 3417-3424, 1987.

155
Oiom

urn

126.0 -i

125.5

r- 60.0

Lounch NA 0.54

Collect NA 0.58 Q Image— shearing values

- 59.5

125.0 -1 - n

.22

I .26

I .30

I .34 .38 "T-

.42

Diameter vs Blocking NA

Fig 1

1

.46

Blocking NA

59.0

Oiom um

125.0 ^

124.8

124.6

124.4

124.2 -

124.0 -

Lounch NA .85

Collect NA .83 Block NA .47

Aperture diameter vs Focus Offset

Fig 2

Focus um

Focus um

Dry cell Launch NA . Block NA

Aperture edge response

Fig 3

-0.8

-1.7

Wet cell Lounch NA

Block NA

n-1.47

Aperture edge response

Fig 4

156
AN ELECTRONIC INDUSTRIES ASSOCIATION INTERLABORATORY

COMPARISON TO RESOLVE DIFFERENCES IN MULTIMODE FIBER

NUMERICAL APERTURE MEASUREMENTS

A. H. Cherin and E. D. Head

AT&T Bell Laboratories

Atlanta, Georgia 30071

D. L. Franzen and M. Young

National Bureau of Standards

Boulder, Colorado 80303

M. Hackert

Corning Glass Works

Corning, New York 14831

INTRODUCTION:

An industry-wide study among members of the Electronic Industries Association (EIA) was

conducted to document and reconcile differences between various numerical aperture (NA) definitions and

measurement methods. Results on twelve multimode graded index fibers indicate that systematic differences

exist among commonly used far-field and index profile techniques. The differences can be explained by a

wavelength dependent factor and the choice of NA definitions. Conversion factors have been developed to relate

the various measurement methods and definitions.

BACKGROUND:

The NA of a graded index multimode fiber is an important parameter that describes a fiber's light

gathering ability. It is used to predict launching efficiency, joint loss at splices, and micro/macrobending

performance. The maximum theoretical NA is a graded index multimode fiber is defined as:

where 9m is the largest meridional ray angle, in air, that will be guided by the fiber. In terms of the fiber's index

profile,

NA, = sinGm (1)

NA, = Vn? - n\ (2)

or

(3)

where

157
H\ — ni

A = for A«l

/ii is the maximum refractive index of the core and n 2

is the refractive index of the cladding. NA can be

determined from a far-field radiation pattern measurement on a short length of fiber. The far-field method is

described in EIA Fiber Optic Test Procedure (FOTP) 47. Using this test procedure, the far-field radiation

intensity pattern, 7(8), of a fiber is acquired and the NAff is defined as the sine of the half-angle where the

intensity has decreased to 5% of its maximum value. For a given wavelength, the relationship between the 5%

far-field numerical aperture, NAff, and the maximum theoretical numerical aperture, NA, is given by: fl]

7(0)

or

7<e) =0.05 = NAj

(4)

NAff = 0.975 NA, (5)

Standards groups are aware that differences exist between NAff and NA, determined by index profile

measurements. Methods currently in use for determining a fiber's refractive index and refracted near-field, RNF,

(EIA FOTP 44). In addition to the difference that is accounted for in equation 5, the different test methods are

commonly implemented at different wavelengths, and NA is a function of wavelength. Most far-field

measurements are made using an incandescent lamp and interference filter at a wavelength of 850 nm; the NAti

obtained from transverse interferometry is measured using a 540 nm line from a mercury lamp; and the NA^

obtained from refracted near-field measurements is measured with a 633 nm HeNe laser. In this work, the

dependence of NA upon wavelength has been measured and a wavelength dependent factor relating the difference

measurement techniques is obtained.

INTERLABORATORY COMPARISON AND RESULTS:

An industry-wide round robin comparison was conducted among nine U.S. and Canadian laboratories

to document the differences among the various methods. Twelve graded index multimode fibers designated A

through L were studied in the comparisons. The most popular core sizes were used—five 50, four 62, one 85

and two 100 micrometer diameter fibers. NAff was determined using the recommendations of FOTP 47 while

NArj and NA* were determined by FOTP 44 and 29 using equations (2) or (3). Average results from the

participants are given in Tables I and II. Nine laboratories submitted far-field results at 850 nm; six submitted

158
refracted near-field results at 633 nm; and one laboratory submitted transverse interferometry results at 540 nm.

Using the far-field method, the wavelength dependence of NA, in a range from 520 to 1550 nm, was also

obtained for several fibers and is shown in Figure 1. The NA difference between 850 nm (NAff wavelength) and

540 nm (//A„ wavelength) is 2.7%, while the difference between 850 nm and 633 nm (NA^ wavelength) is

1.3%.

CONCLUSIONS:

The differences existing in the industry among various NA measurement methods may now be

explained: Measurement wavelength and the definitions applied to the curves account for most of the offsets.

Correction factors may be applied to the properly calibrated RNF and TI numerical aperture measurements to

give the same results as far-field measurements. To obtain 850 nm NA's, 633 nm RNF data should be reduced

by approximately 4% while 540 nm TI data should be reduced by approximately 5%. The relatively large offsets

observed in the the interlaboratory comparison for RNF measurements cannot be explained; perhaps better

methods for calibration are needed.

REFERENCES

1. A. H. Cherin, "An Introduction to Optical Fibers," McGraw-Hill, 1983, p.135.

159
TABLE I

NAfi and NAtl Results

Fiber;

Dia.,/jm NAff SD(( ,% NA,j

% Difference, NA(

1 |-NA|( „ 1An X 1UU

A, 62 0.285 1.4 0.298 4.4

B, 62 0.285 1.4 0.296 r, i 3.7

C, 50 0.226 1.3 0.238 5.0

D, 50 0.226 1.3 0.237 4.6

E, 100 0.291 1.4

F, 50 0.206 1.9 0.213 3.3

G, 62 0.279 1.4 0.289 3.5

H, 85 0.268 1.9 0.284 5.6

I, 100 0.305 1.6

J, 62 0.286 1.4 0.300 4.7

K, 50 0.228 1.3 0.240 5.0

L, 50 0.188 1.6 0.204 7.8

Average 1.5 4.8

Number ol Participants 9 1

TABLE II

NAff and NAfn( Results

Fiber

i

NAff NArn, SDmf> %

% Difference, N„,m - NAii

' xlOO

A 0.285 0.307 5.9 » i B 0.285 0.309 7.4 7.8

C 0.226 0.239 6.7 5.4

D 0.226 0.240 6.3 5.8

E 0.291 0.321 7.8 9.3

F 0.206 0.223 4.9 7.6

G 0.279 0.301 6.3 7.3

H 0.268 0.293 7.5 8.5

I 0.305 0.340 7.4 10.3

J 0.286 0.307 6.8 6.8

K 0.228 0.246 6.1 7.3

L 0.188 0.203 5.4 7.4

Average 6.5 7.6

Number ol Participants 9 6

NA vs. Wavelength

i i i I

i—i—i—i—pi—m—i—r~i—^—i—^"

600

i i i i i i i ii i i y

700 800

Wavelength, nm

900 1300 1550

FIGURE 1

160
Novel Intcrferometric Dispersion Measurement System

A. O. Garg, M. A. Brandtner and C. K. Eoll

Siecor Corporation

489 Siecor Park (RD)

Hickory, NC 28603

Abstract

Several papers (including [1,2]) have been presented on the use of interferometry for

determining the dispersion of various fibers at wavelengths beyond a 1000 nm. Most of

the methods involve critical alignment procedures, during initial setup and at the beginning of each measurement, along with a considerable loss of transmitted power in the

optical paths leading to and from the sample and reference fibers. Typically, such a system

also generates a relatively noisy interferogram, so that its midpoint can be difficult to

pinpoint. The midpoints of the interferograms for a set of wavelengths are used to find

relative group delays which lead to dispersion values.

The system described here is comparatively easy to assemble; relatively few components

require critical alignment procedures, and these procedures are needed only when the

equipment is first set up. None of the components require alignment during normal dispersion measurements. The system also gives smooth interferogram envelopes whereby the

midpoints can be accurately determined with a high degree of repeatability.

Dispersion Measurement System

Two 2x2 single-mode fiber optic couplers (optimized for 1300 nm and 1550 nm) and

precision capillary units are used in a system with rigidly attached and encased optics. As

a result, a minimum amount of careful alignment is needed to set up the interferometric

system. Because capillaries and couplers are used in place of traditional beam splitters,

postioners and lenses for coupling light to and from fibers and sources, no components

need alignment during normal measurements.

The system (Figures 1 and 2) uses a tungsten-halogen light source directly attached to the

monochromator. All four ends of each of the couplers (CI and C2> have been connectorizod. The polished end faces ensure optimum coupling efficiencies. The output of the

monochromator is fed into one arm (11) of a 2x2 coupler (CI I. The output of the monochromator consists of a semi-precision unit housing the coupler end (11). The second input

end (12) of the coupler (CI) is attached to a He-Ne laser which is used for alignment

purposes during initial setup. (It is otherwise switched off). The output ends (01 and 02)

of the coupler (CI) are fed through precision capillary units to sample and reference fibers,

respectively. The sample fiber is directly coupled to the input end (13) of coupler (C2) using

another precision capillary unit. The output end of the reference fiber is mounted in a

161
stopper motor controlled stage. Using lenses, the beam from the reference fiber is eollimated and then focused on the input end (14) of the second coupler (C2). The output end

(03) of the second coupler (C2( is attached to a precision detector assembly housing an

InGaAsP photodiode. The other output end (04) of this coupler (C2) is mounted in another

detector assembly housing a silicon photodiode which detects the output of the He-Ne

laser.

A smooth interferogram envelope (Figure 3) is achieved by use of a small speaker, operating at low frequency, and loosely attached to the input arm (14) of the coupler <C2). This causes the optical path length in the input arm (14) to vary in an oscillatory fashion. As a

result, for each position of the stepper motor controlled stage, a small section of the

interferogram is repeatedly traversed so that the detection electronics can find a point on

the relatively noise-free rms representation of the interferogram envelope. Similar techniques have often been used in differential interferometry (3,4,5].

In calculating dispersion, the raw relative delay data is differentiated numerically - no

curve-fitting is used with the delay data. Instead, curve-fitting is used with the dispersion

values. When this fit is restricted to a wavelength range of 40 nm or less, approximately

centered on the wavelength of zero dispersion, a straight line is a good fitting function [6].

Measurements

Measurements were made on standard and dispersion-shifted fibers. Typical results are shown in Figures 4 and 5. These particular results have been confirmed by means of

standard interferometric equipment as well as a Raman laser system [7]. The total possible error in the wavelength of zero dispersion provided by the new system appears to be less

than +/- 0.5 nm.

The new system permits relatively fast interferometric measurements. A typical set of 200

measurements, involving 10 wavelengths and 20 automated steps per wavelength, provides the zero dispersion wavelength of a fiber in less than 12 minutes.

Conclusion

An interferometric dispersion measurement system which requires no critical alignment

procedure during measurements has been developed. Results are highly repeatable, and

the accuracy is good. The system has advantages in comparison to other interferometric

dispersion measurement systems.

162
References

1) L. G. Cohen and J. Stone, "Interferometric Measurements of Minimum Dispersion

Spectra in Short Lengths of Single-Mode Fibre," Electronics Letters, Vol. 18, p. 564,

1982.

2) F. M. Sears, L. G. Cohen and J. Stone, "Interferometric Measurements of DispersionSpectra Variations in a Single-Mode Fiber", Journal of Lightwave Technology, Vol.

LT-2, p. 181, 1984.

3) A. O. Garg and R. O. Claus, "Optical Fiber Path Differential Wideband

Interferometer", Acoustical Society of America (Chicago, IL), April 1982.

4) A. O. Garg, "Application of Optical Fibers to Wideband Differential Interferometry

and Measurements of Pulsed Waves in Liquids", Master's Thesis, Virginia Tech.

(Blacksburg, VA), July 1982.

5) R. O. Claus and A. O. Garg. "Measurement of Pulsed Ultrasonic Waves in Liquid by

Wideband Differential Interferometry", Acoustical Society of America (Orlando, FL),

November 1982.

6) A. J. Barlow, "Techniques for the Calibration of Chromatic Dispersion Measuring

Instruments", SPIE Vol. 841, Fiber Optic Network and Coherent Technology in Fiber

Optic Systems II, p. 266, 1987.

7) C. Eoll, W. Lieber, M. Loch, H. Etzkorn and W. Heinlein, "Advances in the

Dispersion Characterization of Different Single-Mode Fibers," Technical Digest of the

12th European Conference on Optical Communication, Barcelona, Vol. 1, p. 71, 1986.

+ Voltmeter

Fig. 1. Schematic Diagram of new interferometric

dispersion measurement system

163
Source

Oettcto.

RMS fl»prvMnltflOfl

1nl*ft«regram Enwlop*

Inttrltrogra/n

Fig. 2. Schematic Diagram of delay

times along the main optical paths

Fig. 3. Idealized interf erogram

GROUP DELRY TIME SMF DISPERSION

3.00

^ 2.00

n

CL

o> I .00

— 0.00

0a

•i . 00

1200

*

6

C

W w

1300 MG0 1500 Wavelength Cnm]

10.00

5.00 h

a. 0. 00

'si -5.00

_ - 10. 00

1300 1310 1320 Wavelength Cnm]

Fig. 4. Group delay and dispersion of standard matched-cladding fiber

GROUP DELHY TIME SMF DISPERSION

O.0O

E -1.00

°- i- -2.00 H

-3.00 -

O -4.00 I-

-5.00

1200

E

10.00

5 . CO

- I-T—I "T T T l

1300 MOO 1500 Wavelenath [nm]

+ + * J

1SC0

1553.2 nm

.3744 ps/lnm^frUm)

a 3 . 00

n -5.00

—10. 00

: 520 1530 1540 Wavelength Cnm]

1570

Fig. 5. Group delay and dispersion of dispersion shifted fiber

164
REFRACTIVE INDEX PROFILING - STATE OF THE ART.

K. W. RAINE, J. G. N. BAINES, D. E. PUTLAND.

(National Physical Laboratory, Teddington, Middlesex, U.K.)

INTRODUCTION.

Numerous techniques exist for obtaining the refractive index

profile of optical fibres [1,2]; of these refracted near field

(RNF) is the most commonly used. However, recent intercomparisons

[3.^ i 5] indicate that accurate calibration can be a problem. In

this "state of the art" paper we discuss mechanisms which can lead

to inaccurate RNF measurements, how these effects can be minimised,

and how axial interferometric measurements can be made with an uncertainty of ±1% and be used to calibrate RNF instruments

using a multi-step index fibre. A comparison of measurements made

by RNF and axial interferometry on several fibres is presented

along with some conclusions drawn from recent intercomparisons.

RECENT REFRACTIVE INDEX PROFILE INTERCOMPARISONS.

At least three intercomparisons have been completed in the

last year, these having been organised by Dr. P. Di Vita through the

European Cooperation in Science and Technology (COST) consortium,

by Ir. M. Hutjens of Philips at Eindhoven (Holland), and by Drs.

D. Franzen and M. Young of NBS at Boulder (USA). In the latter

intercomparison the object was the accurate determination of

numerical aperture (NA) from the refractive index profile. The

results of these intercomparisons are due to be published shortly

and cannot be discussed in detail here. However, the general

indications were as follows;

In the intercomparison organised by Philips there were 8

participants who measured a multi-step index fibre with a depressed cladding (mean <5n=-0.006) and a largest step of about

0.018 for which measurements varied between 0.015 and 0.021. The

results tended to group with particular makes of instrument. The

depressed cladding caused vignetting in some instruments and was responsible for the lowest measured values.

The COST intercomparison involved 7 participants measuring 6

monomode fibres having maximum refractive index differences between

0.005 and 0.02; some had simple step index profiles and others

depressed and/or raised annuli. The overall standard deviation

of the results, which did not fit well to a normal distribution,

was ±20%.

Finally the NBS intercomparison which involved 11 participants

measuring 12 multimode fibres with maximum index differences

between 0.015 and 0.030. The NA determined from RNF profiles

varied by ±10% which corresponds to ±20$ in peak index

difference. Although some of this variability could be attributed

to interpretation of index spikes and ripples, the systematic

differences between laboratories indicated that calibration and

measurement problems were an important factor [3].

THE USE OF AXIAL INTERFEROMETRY FOR ACCURATE CALIBRATION.

The main disadvantages of axial interferometry are the

difficult and laborious preparation of the thin fibre slices and

the poor spatial resolution of the technique. These make it

unsuitable for rapid turn-around measurements and its use for

measuring monomode fibres is limited. It does however measure

refractive index differences in terms of the wavelength of light,

providing measurements traceable to national standards and, with

care, uncertainties of ±0.0001 in a refractive index difference

of 0.015 on relatively large and uniform areas of refractive index

165
can be obtained. The technique therefore seems ideally suited to

providing traceable calibrations, via a transfer standard fibre,

for RNF instruments.

In implementing this method an essentially common path

interference microscope is preferred; that is the ray paths

through the imaging optics, sample and air paths should be the same

for the two interfering light beams. This minimises the effects of

aberrations, mechanical instability and air turbulence required

when optical path differences (OPD) as small as 5nm are to be

determined

By placing a Mach-Zehnder interferometer between the objective

and the eyepiece, the Zeiss Interphako interference microscope

achieves common path operation throughout the imaging optics

producing excellent stability and phase uniformity over the field

of view. Operation of the interferometer allows a variable degree

of shear to be introduced between the images formed by the

eyepiece. Thus the fibre core of one image can be superimposed on the cladding of the second image, with cladding against cladding in an adjacent part of the field of view to provide a reference level.

With a lOOum thick fibre slice the combination of an objective

with an NA between 0.5 and 0.8 and a condenser NA of 0.3 provides

an optimum balance between magnification, resolution (about 2um)

and image brightness.

The interferometer is adjusted so that in the absence of any

phase objects the image field is uniform in intensity. A phase

compensator, which changes the path length of one arm of the

interferometer, can be adjusted so that the intensity is minimised

or equalised across an imaged edge of a half shade plate. The

phase plate is situated at an intermediate image plane and provides

a more precise setting than can be obtained from the interference

minimum. Further improvement and ease of operation is achieved by

using a TV system and displaying the intensity profile alongside

the normal image. With a 100pm thick slice this technique can give a single reading uncertainty equivalent to ±0.00005 in

refractive index [6]. Using a phase compensator to measure directly

the optical path removes uncertainties that can arise in

interpreting the intensity distributions of the interferograms but

makes measurements of a complete profile tedious. In addition the

setting is less certain when made on a sloping intensity profile

produced by regions of non-uniform refractive index. However,

perhaps more importantly in this context, effort can be directed

towards averaging the effects of slice imperfections which have

been troublesome in previous work [7]. Measurements can be made

with the image of the core sheared sequentially into the four

quadrants of the cladding and repeated with randomly chosen

starting positions. Possible errors due to imperfections in the

interferometer are overcome by moving the sample stage so that the

readings are taken at the same place in the image plane. Despite

averaging the slices must be of good quality to produce an acceptable standard error of the mean. The quality of the fibre slices can be checked by interference

microscopy and self-Fizeau fringes which can be seen with the

unaided eye in the capillary section using a He-Ne laser or low

pressure mercury lamp. Once assured that the parallelism of the

capillary is satisfactory the interference microscope can be used

to shear the image of the fibre into the capillary to assess the

quality of the fibre slice. As a second test the end faces of the

fibres can be viewed with a microscope incorporating a Michelson

interferometer objective. This will show end face flatness and any

tilt of the fibre slice with respect to the capillary. Typically

thickness variations less than O.lum can be achieved and

deviations from flatness less than 0.05um. The overall

uncertainty in the OPD measurement due to slice imperfections is

thought unlikely to exceed ±0.5% when the averaging measurement

166
procedure described above is adopted.

The thickness of the fibre slice can be measured using the

z-axis measurement facility on the Vickers Fibercheck microscope.

Precision focussing on the upper and lower surface of the slice is

achieved by aligning two halves of a graticule image and an electronic readout of the displacement is given with an uncertainty

of ±0.5um (2a). The instrument is calibrated using a thin silica slice whose thickness is known from interferometric

measurements. The effects of slice imperfections are averaged as in the phase measurement which also reduces the random uncertainty.

The overall uncertainty of the thickness measurement is considered

to be ±0.2um.

In summary, the overall uncertainty in the refractive index

difference is systematic and can be estimated to be ±0.7% or ±0.0001 in an index difference of 0.015 (la).

REFRACTED NEAR FIELD.

The difficulties in making accurate RNF measurements are due

to uncertainty in the calibration techniques and to making accurate

radiometric measurements particularly when there are angle of

incidence changes in the intensity distribution.

The most common methods of calibration involve using the

oil-cladding index difference, a fibre of known NA and moving the

blocking disk. Oils have large temperature coefficients of

refractive index requiring the temperature to be known to better

than 0.3°C and the refractive index of the cladding can be

several parts in the 4th decimal place higher than that of bulk

silica [6]. Use of a fibre of known NA relies on definitions, and on the interpretation of imperfections, ripple, spikes and dips in

the refractive index profile. For the purpose of traceable

calibration and obtaining the fullest understanding and confidence

of measurement techniques it is more satisfactory if the value of

the NA calculated from the profile agrees with that measured by

far-field and vice versa. A direct refractive index calibration

using a fibre measured by interferometry therefore seems to be the

preferrable choice. It would provide traceability for and

confidence in other methods perhaps preferred for convenience by

other laboratories.

SOME RNF MEASUREMENT PROBLEMS.

Movement of the blocking disc only simulates movement of the cone of rays. This is subject to errors if the transmissivity of the

collecting optics significantly varies with angle of incidence

(considered by Stewart as a potential source of error [1]). This

effect can be seen if an extra blocking disc is placed close behind

the launch objective so that a hollow cone of light is focussed

on to the end of the fibre. The size of this disc should just

prevent the innermost rays intercepting the instruments blocking

disc. Thus when the focussed spot is scanned across the fibre the

output from the detector should be constant. Figure la shows

instead an inverted profile which has the effect of reducing the

profile height and the height of the oil-cladding step by at least

0.0005 in index. Figure lb shows the strong polarization

dependence and further emphasises the need to use circularly

polarized light or light plane polarized at 45° to the scan

direction. These error curves can then be used to produce

correction curves as a function of index difference; the curve for

our instrument is shown in figure 2. Alternatively the collecting

optics should be designed to ensure the angles of incidence at

air-glass interfaces are less than 30° or correcting optical

coatings used, or, in the case of ellipsoidal collectors, an eccentricity of 0.71 ensures that angles of incidence are within a

167
few degrees of 45°; small changes in angle of incidence can then be made to have an insignificant effect.

We have also observed with our instrument that repeatable and

consistent measurements are only obtained from the average of

orthogonal scans and that the oil index level often needs to be

determined a distance greater than 30um from the oil-cladding

interface. These effects seem to be due to the changes in annular

light intensity distribution during a scan. The detector needs to

be insensitive to this and should not be driven non^linear by high

localised intensities which must not exceed 70uW/mm . Thus to avoid possible detector non-linearities and position

sensitivity of the detector a blurred spot is preferrable to a sharply focussed point or a well defined annulus. Even so, the

badly focussed spot size will vary and vignetting at the detector

must always be considered. These difficulties can be overcome by

placing an integrating sphere in front of the detector, the

focussed light now falling onto a diffusing baffle in the centre of

the sphere. This effectively desensitises the detector to annular

and angular distribution variations in the incident light. We have

found its use beneficial to the production of level symmetrical

profiles on step fibres and in reducing scan direction dependence.

The reduced detection efficiency did not lead to signal-to-noise

problems

In an effort to ensure some measurement repeatability our RNF

profiler was assessed and modified in light of the points discussed

above. The movable blocking disc was repositioned so that there was agreement with the axial interferometric results on the

calibration fibre (whose RNF profile is shown in figure 3) after

the angle of incidence corrections had been applied. Comparisons

between RNF and axial interferometry were then made on several

other fibres, one having a step index and the others parabolic

index profiles with no, or very small, on-axis index depressions.

The calibration and refractive index of the oil were used to

calculate the refractive index of the cladding which was also

determined by phase contrast microscopy [8]. The NA of multimode

fibres was evaluated from the smoothed curve fit profiles [9] and

compared with the Far-Field values from the NBS intercomparison.

The results are shown in table 1. RNF measurements are the average

of at least four sets of orthogonal scans, which also reduced the

random uncertainty to ±0.0002 (2a). The refractive index

of the oil was measured at 20°C on an Abbe refractometer

and the oil temperature by a platinum resistance thermometer

adjacent to the cell; during RNF measurements it was necessary to

minimise temperature gradients by allowing the instrument

temperature to stabilise (for a few hours) and the oil cell, after

loading, to come to thermal equilibrium with the instrument.

Table 1 shows that sufficiently accurate results can be

obtained by RNF measurements. Calibration via interferometry and

the application of the described correction curve leads to a ±1% agreement on index difference over a range of fibres, a ±2% agreement on the cladding refractive indices and a +1%

difference (average of 12 fibres) in the NA determinations.

CONCLUSION.

Accurate measurement and calibration of RNF instruments can be

achieved with uncertainty of ±1%. Accurate operation may

require in some cases that additional attention be paid to the

collecting optics and detector, angle of incidence error curves produced and scan direction dependencies investigated. A direct

refractive index calibration, traceable to national standards of

length is possible using multi-step index fibres measured by axial

interferometry. The fibre whose profile is shown in figure 3 is

now available with a certificate of calibration. With accurate

168
refractive index and geometry calibration (also under investigation

at NPL) an RNF instrument can provide measurements of profile

height, profile shape, all of the geometrical parameters and the

NA.

Table 1

Fibre type

Multi-step

Single step

Multimode

Index difference

RNF AI

Cladding index

RNF PCM

NA

RNF* FF

0.0184

0.0147

0.0037

0.0184

0.0148

0.0036

0.0162 0.0163

0.0310

0.0292

0.0160

0.0196

0.0307

0.0291

0.0160

1.4577 1.4576

1.4576 1.4577

1.4575 1.4575

0.293 0.286

0.281 0.279

0.209 0.206

0.228 0.226

notes AI = axial interferometry

PCM = phase contrast microscopy

FF = far- field

* = corrected to 85O nm wavelength

REFERENCES.

[1] STEWART, W. J.: "Optical fibre and preform profiling

techniques.", IEEE J. Quantum Electron. QE-18, 1451-1466, 1982.

[2] MARCUSE, D. : "Principles of optical fiber measurement. " .Academic

Press, Chapter 4, 1981.

[3] FRANZEN, D.L., YOUNG, M. , CHERIN, A.H. , HEAD, E.D.,

HACKERT, M., RAINE, K.W. and BAINES, J.G.N. ,: "Numerical aperture

of multimode fibres by several methods: Resolving differences.",

Submitted to J. Lightwave Technology.

[4] Di Vita, P.: "Report on COST 217 intercomparisons on circularly

symmetric single mode fibres.", To be published.

[5] HUTJENS, M. -."Round-robin results of fibre refractive index

differences measured by the refracted near field (RNF) method.",

To be published in J. Opt. Comm., 9, 4, 1988.

[6] RAINE, K.W., BAINES, J.G.N. , KING, R.J. : "Comparison of

refractive index measurements of optical fibes by three methods.",

IEE Proc. J., 135, 190-195, 1988.

[7] W0NSIEWICZ, B.C., FRENCH, W.G., LAZAY, P.D. and SIMPSON, J.R. ,

"Automatic analysis of interferograms : optical waveguide refractive

index profiles.", Appl. Opt., 15, 1048-1052, 1976.

[8] LOCKE, J., UNDERHILL, M. : "Automatic refractive index

measurement of glass particles.", Forensic Science International,

27, 247-260, 1985.

[9] FORBES, A.: "Fitting a generalized parabola to data", NPL Report

DITC 96/87, September 1987.

ACKNOWLEDGEMENT.

The authors would like to thank Ir. M. Hutjens and Dr. P. Di Vita

for permission to discuss the results of the intercomparisons

ahead of publication.

169

RECONSTRUCTED REFRACTIVE INDEX PROFILE OF OPTICAL

FIBER BY REFRACTED RAYS

Nobuo AMANO and Ken-ichi NODA

Faculty of Technology

Tokyo University of Agriculture and Technology

Koganei-shi, Tokyo, 184, Japan, telephone +81 423 81 4221

New microscopic computer tomography measurement method of optical fiber

refractive index profile has been reported by the authors elsewhere [1 ]. In

the first half of this paper, this microscopic CT technology is applied to

polarity-maintaining optical fiber refractive index profile measurement. It

is difficult to measure it by conventional method because of masking effect of

the stress-giving area. In such a case, effectiveness of this method is

demonstrated by showing experimental result.

In our previous paper, error estimation of this method which may be

produced in the process of the ray to index-strip-integral transformation was

not enough. So in the latter half the error estimation of this method is made

through a computer simulation for a given input refractive index distribution.

In the computer X-ray tomography, an X-ray flux passes through an object,

such as a human brain, and a received projection represents a strip integral

of the absorption in the object. Unlike the X-ray, in most optical material,

light does not suffer any absorption but is refracted due to the refractive

index non-uniformity [2-3]. Therefore, a received projection represents a ray

flux density distribution due to refraction. New and improved signal processing algorithm is needed. It is that the ray flux density distribution is

transformed to the strip integral of the refractive index along the projection

path, as described in Ref. 1.

This method has advantages that there is no need to prepare special

samples, and only ordinary microscope and TV camera are used and is a nondestructive measurement. This is especially significant for nonaxisymmetrical

distributed refractive index profile measurement. Effectiveness of this

measuring system has been demonstrated by showing an example of measured

171
result on a dual core fiber who has two single mode cores in one cladding [ 1 ]

This time, elliptic jacket fiber was used as measured material as an example of the polarity-maintaining fibers. This fiber has nonaxisymmetrically

located stress-giving portion near the core to produce large birefringence in

the core. Although residual stress distribution in this kind of optical fiber

has been measured using CT algorithm [A], refractive index profile measurement

has not yet been reported. It is not possible to measure, using observed data

from only one side of the fiber. Moreover, in general, stress-giving portion

has lower refractive index than that of the cladding material and has large

cross sectional area. So, it is difficult to measured the core profile by

usual optical observing method, because non-uniform refractive index region

distorts the image of the core when observed from the outside of the fiber.

In our measurement, the fiber immersed in the matching liquid is observed

with a microscope and television camera, and 480 times 480 pixels of eight

bits multilevel light intensity data are stored in a frame memory. The obtained data are averaged in the fiber longitudinal direction, to minimize

noise due to dust, and other causes. This observation and averaging are

repeated rotating the fiber by t /n at a time, where n is a projection number

of CT. Then n intensity distribution data items are obtained in w radians of

rotation, around the axis. The observed data are transformed to the index

strip integral. At each rotation angle of view, the observation and calculations are performed. After repeating these procedures, the obtained all

projections are space-filtered and back-projected to reconstruct the refractive index two dimensional profile [5]. n=100 was selected in the most

measurements. An example of three dimensional plot of reconstructed refractive index profile is shown in Fig. 1 . The measured profile along the major

and minor axes are shown in Fig. 2. In these figures, a higher refractive

index region of a small core is clearly demonstrated.

A computer-simulated experiments has been performed to estimate the error

of the microscopic CT algorithm. Fig. 3 shows an example of simulation

172
result. In Fig. 4, an input refractive index profile for the simulation is

given, by a broken curve, which has five concentric circle areas. There are

four levels of refractive indexes, including a low level of equivalent of

stress-giving portion. Fiber radius of 60 fim, core radius of 5 n m, observing

area diameter of 150 urn are assumed. Reconstructed index distribution along

the fiber diameter line is shown in Fig. 4 by a solid curve, and is compared

with the input data. The reason why axisymmetric index distribution is

selected as an input, is merely to make the error estimation easy.

Ray trace simulation is done as follows; white light ray flux enters into

a hypothetical optical fiber having a given input index distribution, and

light intensity distribution due to the illuminated light is calculated. The

refraction and reflection at the boundary cylinders between neighboring concentric circle regions are obtained by Snell's law. The reflection and transmission coefficient at the boundary cylinders are obtained by Fresnel's law.

The scattering at the boundary cylinders is ignored. Pixel number at the observing plane is 300 dots. The ray flux number is assumed to be 100 per a

pixel. 30 000 flux in total are used in this simulation. The reconstruction

is made by this method using these data.

In this simulation, variance of the error is 0.57. In this value, ray

trace simulation error is also included. There are a little error of the artifact accompanying CT. Reconstructed profile by the simulation is very

similar to that of the given profile, except near the stepwise index change

portions. If the refractive index change is rapid, ray crossovers may cause a

distortion of the reconstructed shape. But it is not so important because

practical fiber has a relaxed index change as shown in the experimental

result. The obtained variance is mainly attributed to the given stepwise index change.

In conclusion, an elliptic jacket optical fiber having nonaxisymmetrically distributed refractive index, is measured by the microscopic

CT, and index profile of small core surrounded by a large area of lower index

173
stress-giving portion is successively demonstrated. The error estimation

through a computer simulation is made.

REFERENCES

[1] K. Toga, et al, Jour. Lightwave Techn. 6, 1, pp. 73-79, Jan. 1988.

[2] D. Marcuse, Appl. Opt., vol. 18, no. 1, pp. 9-13, Jan. 1, 1979.

[3] D. Marcuse, and H. M. Presby, Appl. Opt., vol. 18, no. 1, pp. 14-22, Jan.

1, 1979.

[4] T. Abe, et al, Electron. Lett., vol.21, no.1, pp. 4-5, 1985.

[5] L. A. Shepp and B. F. Logan, IEEE Trans. Nucl. Sci., vol. NS-21 , pp. 21-

43, June 1974.

MINOR AXIS

MAJOR AXIS

Fig. 1 3D plot of reconstructed index profile of the elliptic jacket fiber.

-50 -25 0 25 50

FIBER RADIUS [ M m]

Fig. 2 Major and minor axial direction distributions of the

elliptic jacket fiber refractive

index.

Fig. 3 3D plot of reconstructed refractive index profile by a simulation.

-50 -25 0 25 50

Fig. 4 Input (broken curve) and

reconstructed (solid curve) index

distributions along the diameter

direction.

174
CALIBRATION OF OPTICAL FIBER PREFORM REFRACTIVE INDEX MEASUREMENT

D. A. Svendsen , G.Beresford and D.Marcus-Hanks

York Technology Limited, York House, School Lane,

Chandler's Ford, Hants S05 3DS, United Kingdom.

Introduction

The measurement of the retractive index of optical fiber preforms by

transverse illumination can achieve repeatabi 1 i ti es of better than 100 ppm in

refractive index. For example, a measurement of a graded-index niultimode fibre

preform gave a core index difference of 0.01265, with a one standard deviation of

0.00008, based on 9 groups each of 9 measurements made over a period of several

days. To fully realise this performance and provide the basis for substantial

improvement, a calibration technique is required that offers of the order of 0.1'/.

of typical index differences. The preform analyser obtains the index profile from

the deflection function which is the change in refracted angle as a function of

preform radiusCl]. It is proposed that the angle measurement is calibrated rather

than the index difference directly. Existing calibration techniques rely on

measuring index differences known by some independent reference method; for

example the i nterf erometri c measurement of a slice of the preform[2]. However,

such techniques are generally destructive, unsuitable for routine use and may not

give the required accuracy. The three different approaches to angle calibration

described offer both convenience and high accuracy, and will be traceable to

absolute standards using conventional index and dimensional measurements.

Calibration lens in air

A typical preform analyser consists of a light source, the preform immersed in

index matching oil inside a cell, and a means of measuring the refraction angle as

175
The gradient at the centre of the tube is thus:

d Q(y) ^_ -2Nlf/ N2 - Ni l + / N3 - N2 |Nl]

y=0 II Rl N2 / I R2 N3 / N2 (iv)

It is possible to correct for the absolute index of the matching oil, and

hence its temperature sensitivity, so that the deflection function depends only

upon the silica/air interface. Figure 2 shows some typical results for the

measured gradient for a 42mm diameter, 2mm thick wall, fused silica tube, both

with and without temperature correction. Repeat ab i 1 i t i es of well under 1 part in

a 1000 have been achieved using standard tubes. The circularity and wall

thickness variations of these tubes are being investigated to ensure that the

required accuracy may be achieved by computation using expression (iv).

Pr i sms-i n-cel

A less convenient, but potentially more accurate, alternative to the tube is a

right-angle prism immersed in the cell. With the light source at tilt angle A to

the normal of the hypotenuse face, the deflection function is given by:

0 = -B + ARCSIN N2 SIN[A + B - ARCSIn/nJ_ SIN(A) 1

Nl \ \ N2 (v)

where the terms are defined in Figure 3. With silica index matching oil (Nl=l .458)

and a crown glass pr i sm <N2= 1 . 51 96) , the difference in refraction angles between

the two exit surfaces is about 7 degrees (Figure 4) with a sensitivity to tilt of

the prism of about 0. 0015deg/deg A) and to prism angle error of .09deg/deg

(^0/^B). Thus a prism with angles accurate to better than larc. minute and tilt

less than 3 degrees will meet the 0.17. required accuracy if the refractive index

is known to 50 ppm. The absolute index of the oil and its temperature sensitivity

may be eliminated by utilising an adjacent fused silica reference prism. Although

the prism only gives one calibration angle unlike the other two methods, it has

the advantage that it does not require the cell movement stage to be calibrated.

In addition, the significance of strain in all three calibration samples is

being studied to ensure that the stress optic effectC3] does not degrade the

absolute accuracy.

176
the preform tra ver sesC 1 ] . The simplest approach to angle calibration is to

substitute a standard plano-convex lens for the cell and preform. The deflection

function for normal incidence on the plane surface of the lens is then:

0<y) = ARCSIN /N2 A - ARCSIN /A

\HZ RJ

\nj (i)

where: y is the distance from the centre of the lens, R its radius of curvature,

and N2,N3 are the refractive indices of the lens and air, respectively.

The gradient at the centre of the lens is thus:

d 0(yl

dy — Z n2 - N3 \

\ R N3 /

(i i

)

The refractive index is known to better than lOppm and so the gradient

accuracy will be limited by the calibration of the cell movement stage, typically

about 0.17.. Not only may the gradient be used to calibrate the measurement to the

required accuracy, but the known shape of the theoretical deflection function

allows the linearity of the instrument to be checked. Figure 1 compares the

deviation of expression (i) from a linear function with that of an actual

measurement. Short term repeatabi 1 i ti es of 1 part in 3000 and longer term of 1

part in 1000 have been achieved, hence coming close to the target of 0.17.

accuracy.

The objection to this approach is that the cell must be removed which may

introduce errors arising from the absence of the cell windows from the optical

path

Tube-i n-cel

A hollow fused silica tube inside the cell gives a deflection function of the

form:

0 (y) = -2N1 ["ARCSIN/^. \ - ARCSIN/_N1 ^_\+ ARCSIN/nU _y_\- ARCSIN/ni _y_

j \Rlj [n2 Rl/ 1N2 R2/ In3 R2

(i i i

where: y is the distance from the tube axis, Ri,R2 are the outer and inner radii

of the tube, and N1,N2,N3 are the oil, tube and air refractive indices

respectively.

177
ftcknowl edgement

Thanks are due to the Directors of York Technology Limited for permission to

present this paper.

References

CI] W.J.Stewart, "Optical Fiber and Preform Profiling Technology", IEEE

J. Quant. Electron. , 1982, Vol. QE-1B, N010, p 1451-1466.

[23 M.Carratt, C.Reinaudo, J.Auge, "Preform measurements and their industrial

application", 1985, SPIE Proceedings, Vol.584, paper 584-19, p 1 22- 128

C33 G.W.Scherer, "Stress-induced index profile distortion in optical waveguides"

Applied Optics, 1980, Vol.19, Nol2, p2000-2006.

H

I)

V

~0

>

1)

. 04

. 02

-.02

. 04

' ' I

' 1

' I

' '

theory

|'l'(,y measurement

'.l.i.l. i.l.i

0 2 Fioure 1. Lens

1 0

1 . 003

- 1.002

c

v

! . 00 !

z 999 — O compensated

'_ uncompensated

22 23

radius ( mm

)

20 2 1 Figure 2. Tube

temperature (deg C)

A

B

e

cngle of till of prism

angle of exit face

deflection cngle

Figure 3. Print!

-4-2 0 2 4 Figure 4. Prism

entry point on hypotenuse (mm)

178
INTERLABORATORY COMPARISON OF FAR-FIELD METHODS FOR DETERMINING

MODE FIELD DIAMETER USING BOTH GAUSSIAN AND PETERMANN DEFINITIONS

Timothy J. Drapela

National Bureau of Standards

325 Broadway

Boulder, CO 80303

In a previous interlaboratory comparison of measurement methods for mode

field diameter (MFD) of single-mode optical fiber, participants assumed a

Gaussian mode-field distribution [1], Those results indicated good agreement

(0.15 ftm, one standard deviation) among the various methods for dispersion

unshifted fibers at 1300 nm. The Gaussian distribution was a good assumption

since 1300 nm was close to the cut-off wavelength. At 1550 nm, however, the

Gaussian assumption is not valid, and significant offsets (0.7 /xm) were

observed between the one -dimensional far-field scan (FF) method and variable

aperture far- field (VAFF) method. This is of concern, since these two methods

are among the most widely used for routine MFD measurements. Anderson and

Kilmer have shown that the "Petermann 2" definition, based on a second moment,

gives good consistency among various measurement methods [2] [3] . Moreover, in

those situations where mode profiles are not approximately Gaussian, the

Petermann definition gives better prediction of splice loss than does the

Gaussian.

This paper describes an interlaboratory comparison which includes the

Petermann as well as the Gaussian definition and also considers measurements on

dispersion shifted single-mode fibers. Participants in the comparison are

members of the Electronic Industries Association and include most of the major

fiber and cable manufacturers in North America.

Six single-mode fibers representing five different manufacturers were

measured by the participants. Two fibers were dispersion shifted, while the

remaining four were dispersion unshifted.

179
Seven participants used the FF method, in which laser diodes are used as

sources and MFD is determined by acquiring the one -dimensional far- field

radiation pattern. Five participants used the VAFF method, in which MFD is

determined by measuring the relative power passing through a series of farfield apertures. Here, the source generally consists of a tungsten lamp and a

monochromator . Details of the FF and VAFF methods can be found in EIA-FOTP 164

and EIA-FOTP 167, respectively.

Common to all methods were a test fiber length of 2.0 ± 0.2 m, a single

50 mm diameter loop to strip out possible second-order mode power, and, when

necessary, some type of cladding-mode stripper.

Participants closely followed the guidelines of FOTPs 164 and 167. The

FOTPs contain specific curve-fitting instructions for the Gaussian MFDs.

Petermann MFDs were determined from far-field integrals without specification

of a particular curve- fitting routine.

The results of the comparison are grouped into three categories: 1300 nm

dispersion unshifted, 1550 nm dispersion unshifted, and dispersion shifted.

This division seems quite natural when the average offsets between FF and VAFF

methods are examined. For fibers in the same category, the relative offsets

are approximately the same, without exception; relative offsets between FF and

VAFF methods are given in Table 1 (small corrections have been made for

variations in source wavelengths) . The offsets are less for the Petermann

definition for all fiber categories.

For unshifted fibers at 1300 nm, using the Gaussian definition, the results

of the previous interlaboratory comparison were verified. An average standard

deviation of 0.17 /xm was observed, with only a small offset between average

VAFF and FF results. The Petermann results did improve on this though, giving

almost no offset and an average standard deviation of 0.11 /im, Fig. 1.

180
For the same fibers at 1550 nm, substantial offsets were observed between

VAFF and FF results for the Gaussian definition. Petermann results gave

significant improvement, reducing offsets to almost zero and giving an average

standard deviation of 0.11 /im, as opposed to 0.32 for Gaussian, Fig 2.

With dispersion shifted fibers, the Gaussian results gave the largest

offsets of all. Petermann results seemed to improve upon this, but only

slightly. The average standard deviation was 0.5 fxm for Gaussian and 0.35

for Petermann, Fig 3. The Petermann results of one VAFF participant were

consistently within the range of FF Petermann results. We suspect the large

VAFF/FF offsets are related to the small MFDs of dispersion shifted fiber.

Smaller MFDs translate to more power at larger far-field angles. If the VAFF

optics and associated detector do not have sufficiently uniform collection for

large enough NA, then the measured far- field pattern will be truncated; this

translates into a larger measured MFD. We believe, therefore, that the

discrepancies observed for this fiber category have more to do with

experimental apparatus than with the Petermann definition; this is currently

being investigated.

REFERENCES

[1] D. L. Franzen and R. Srivastava, "Determining the Mode-Field Diameter of

Single-Mode Optical Fiber: An Interlaboratory Comparison," J. Lightwave

Technol., vol. LT-3, no. 5, p. 1073, Oct. 1985.

[2] W. T. Anderson, V. Shah, L. Curtis, A. J. Johnson, and J. P. Kilmer, "ModeField Diameter Measurements for Single-Mode Fibers with Non-Gaussian Field

Profiles," J. Lightwave Technol., vol. LT-5, no. 2, p. 211, Feb. 1987.

[3] K. Petermann, "Constraints for Fundamental Mode Spot Size for Broadband

Dispersion-Compensated Single-Mode Fibers," Electron. Lett., vol. 19,

no. 18, p. 712, 1983.

181
Table 1. Average offset values,

in um, between VAFF and FF results

for the three categories of

comparison fibers.

Category

Offset Between VAFF and FF Methods, am

(Ave. VAFF - Ave. FF)

Petermann Gaussian

1300 nm

Dispersion

Unshifted

0.04 0.28

1550 nm

Dispersion

Unshifted

0.05 0.59

1300&lS50nm

Dispersion

Shifted

0.52 0.92

10.0

E

E

Dispersion Unshifted

Fiber 2 ; 1300 nm

_ Petermann

3

9.0

Gaussian _

: +

+ VAFF ; • FF Scan

Fig. 1. Typical result for

dispersion unshifted fibers

at 1300 nm.

Dispersion Unshifted

Fiber 4 ; 1550 nm

_ Petermann

12.0

E

0

E

S 11.5 0

Ll.0

"8

n.o

Gaussian

+

+

+

• • •

• +

+ VAFF ; »FFScan

Fig. 2. Typical result for

dispersion unshifted fibers

at 1550 nm.

Dispers

Fiber 3

on Shifted

1550 nm

8.5

E

=*• 8.0

E

<Q Q

TJ0

.7.5

7.0

Petermann Gaussian

+ VAFF ; «FFScan

Fig. 3. Typical result for

dispersion shifted fibers.

182
IMPROVEMENTS TO THE VAMFF SYSTEM FOR ACCURATELY

MEASURING THE MODE-FIELD DIAMETER OF DISPERSION SHIFTED FIBERS

JERRY R. PARTON, PRODUCT ENGINEERING, MP-RO-2

CORNING GLASS WORKS, CORNING, NEW YORK 14831

Introduction Mode-field diameter (MFD) is an important parameter in the

performance and specification of optical fibers*. This places a requirement

on the industry to have accurate MFD measurement methods so that this

parameter can be controlled and specified. Improved measurement accuracy

will provide improved manufacturing process feedback, while facilitating

better agreement throughout the industry.

2

Improvements to the variable aperture method in the far-field (VAMFF)

MFD measurement system are presented. The principle modification is the

increase of the maximum numerical aperture (NA) of the system by increasing

both the number and size of the apertures. By increasing the maximum NA from

.2 to .4, sufficient power is collected from the tails of the far-field power

distribution of dispersion shifted (DS) fibers to allow a more accurate

measurement of these non-gaussian mode-fields. Without this modification,

the .2 NA system can result in MFD error of up to 7% for some DS fiber

designs. Agreement to within .1 urn is demonstrated between the new "High NA

VAMFF" system and one-dimensional direct far-field scan measurements on DS

3 fibers based on the Petermann II definition of MFD.

Equipment The increased NA is accomplished by a spherical mirror focusing

the light back through the aperture and onto the detector. The fiber and

detector are each offset from the mirror's axis by .85 mm yielding a 1.7 mm

fiber to detector separation (see Figure 1). The aberration introduced by

183
this offset does not negatively impact the system because the 300 urn detector

diameter was appropriately chosen to collect essentially all of the power in

the focused spot.

Vignetting is minimized by the close spacing of the components, and the

remaining vignetting effect is compensated for by a geometric calculation of

an effective aperture size. An additional benefit to using a mirror rather

than a lens system is the absence of chromatic aberration, allowing the

system to be used over a broad range of wavelengths. Twenty-one apertures,

ranging from .03 to .40 NA, are used in the aperture wheel, providing good

coverage between the largest and smallest apertures.

Calculation DS fibers, having more complex index profiles than standard

step-index fibers, also have more complex non-gaussian field distributions.

The Petermann II shape-independent definition of MFD is an accepted

4 definition for the fundamental-mode spot size of DS fibers. Pask has shown

that this definition can be transformed to the far-field, from which point

integration by parts yields the following expression for MFD:

MFD = 2

'

2 /V 2 1

1 2

' ff

where W,

2 = K2 [sin2

9 -(2/P XJ9"1 P(9) sine dsinO)]

ff 1 m m' wo

K = 2n/X

9

m

= half angle of largest aperture

P = power measured through largest aperture

**(9) ~ Power measured through each aperture

The integral is solved numerically. This can be done by trapezoidal

numerical integration, but depending on the form of the integrand and

differential used, the result may vary across integrands by more than one

percent. Recent tests have determined, however, that numerical integration

184
by Simpson's Rule is self-consistent across the same integrands to within

.2%, and is centered between the errors of the trapezoidal methods. Use of

Simpson's Rule can therefore give a slight improvement to agreement between

laboratories.

Results The High NA VAMFF system has been compared to one dimensional

direct far-field scanners in two separate forums. The first was the 1987

CCITT Study Group XV contribution 145 entitled Experimental Data for

Mode-Field Diameter by the Variable Aperture Technique. These results are

summarized in Table 1. The second comparison came from the 1988 NBS/EIA

measurement round robin for MFD. These results are summarized in Table 2.

In general, the MFD bias between the High NA VAMFF system and the various

far-field scanners was less than .1 urn, with a VAMFF measurement standard

deviation of approximately .05 urn.

Conclusions Modifications to the VAMFF system have made it possible to

accurately measure the MFD of DS fibers by this technique. The primary

modification is the use of a spherical mirror to collect power out to .4 NA.

The calculation accuracy is also improved by using Simpson's Rule to carry

out the numerical integration. Experimental results show excellent agreement

between High NA VAMFF and one-dimensional direct far-field scanners.

Therefore, the variable aperture technique, with its relatively short

measurement time, can be used to accurately measure MFD of both standard

step-index fibers, and dispersion shifted fibers.

Acknowledgments The "High NA VAMFF" system was initially designed and

tested by R. A. Modavis and A. B. Bussard of Corning Glass Works.

185
References

1

D. Marcuse, BSTJ, Vol. 56, No. 5, p. 703, 1977.

2

J. Dick et al., OFC Technical Digest, paper WB3, 1984.

3

K. Petermann, Electronics Letters, Vol. 19, No. 18, P. 712, 1983.

A

C. Pask, Electronics Letters, Vol. 20, No. 3, p. 145, 1984.

Ap«rtur«

Figure 1 (not shown to scale)

High NA VAMFF Optical Path

Table 1 - CCITT Study Group XV, 1987

Mode-Field Diameter Contribution 145

Fiber

1

2

Wavelength

(nm)

1550

1537

High NA VAMFF

Mean (urn)

8.338

8.183

Far-Field Scan

Mean (urn)

8.374

8.155

average bias =

Bias

(urn)

-.036

.028

.004 ym

Table 2 - NBS/EIA 1988 Mode-Field Diameter Round Robin

Fiber

3

4

3

4

Wavelength

(nm)

1550

1550

1300

1300

High NA VAMFF

Mean (urn)

7.06

8.19

5.77

6.42

Far-Field Scan

Mean (ym)

7.07

8.34

5.83

6.48

average bias =

Bias

(um)

-.01

-.15

-.06

-.06

-.07 ym

186
ACCURACY OF NODE FIELD DIAMETER MEASUREMENTS IN SINGLE-MODE FIBRES

FROM NEAR-FIELD TECHNIQUE

M. Artiglia, G. Coppa (°), P. Di Vita, M. Potenza

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

ABSTRACT

The accuracy of Transmitted Near-Field Intensity measurements for the

determination of Mode Field Diameters in single mode optical fibres is discussed. The

possible error causes are analyzed, with particular attention to systematic terms, and

an evaluation of their effect is performed.

1. INTRODUCTICN

One of the most important dimensional parameters for the design and the

characterization of single-mode optical fibres is the Mode Field Diameter (MFD).

Various techniques for the measurement of the MFDs have been proposed in the

literature. Among these approaches, the one based on Transmitted Near-Field Intensity

(TNFI) scanning is particularly attractive because it can be used also for geometrical

measurements [1J. Aim of this paper is to carry out a comprehensive analysis, still

lacking, on both the accuracy and the reproducibility of the technique to the end of

providing valuable suggestions for its practical implementation. In particular, the

delicate case of dispersion modified SM fibres is examined for its actual interest.

It is well known that from a TNFI scan both Near-Field (NF) MFD and Far-Field (FF) MFD

can be calculated from the relations [2]:

d^- 2 /2 { J rV(r) r dr / JeV) r dr }

/z

(1)

o o

d - 2 /2 { J E

2

(r) r dr / J (dE(r)/dr)

X

r dr }

/z

(2)

J 6 o

where r is the radial coordinate on the fiber endface and E(r) is the LP01 electric

field. In order to evaluate the accuracy in the MFD determination, it must be pointed

(°) Present address:

POLITECNIOO di TORINO - C.so Duca degli Abruzzi, 24 - Torino (ITALY)

187
out that such a measurement is performed in two steps. First, the TNFI pattern is

acquired, successively the experimental data are numerically processed according to

the defining equations. The TNFI acquisition is affected both by the measurement

noise and by the diffraction effects on the image of the NF. Thus an analysis of the

TNF optical system (supposed to operate in diffraction limited conditions and with a

spatially coherent illumination) is needed in terms of the objective NA and of the

dynamic range of the acquisition system, which are the quality parameters of the

experimental set-up. Furthermore the numerical processing of experimental data can

introduce additional errors that must be taken into account. All of these error

causes will be discussed in the following.

2. ACCURACY OF THE NF MEASUREMENTS

The aim of a TNFI measurement is the determination of the width of a light

intensity distribution, acting on a magnified image of it. The possible causes of

errors are due to: accuracy in the positioning of the scanning motors, accuracy in

the determination of the optical system magnifying power, use of a finite aperture to

collect the TNF local intensity on the image plane (errors due to the finite aperture

of the lens will be discussed in the following Section; errors introduced by improper

stripping of the fibre or by uncorrect lens focusing can be removed in principle and

will not be discussed here). The accuracy of the traslation stages is within 1 //m on

the image plane. Since the lateral magnifying power M is around 50X this corresponds

to an error of + 0.01 /m on the fibre end face. TO determine the magnifying power

M, a precision bar mask can be employed as optical object to calibrate the spatial

scale on the image plane; taking again into account the accuracy of translation

stages, the error on M can be evaluated around + 0.1X, thus contributing a further

+ 0.02 /m on the fibre face. It is interesting to note that the effect of the

finite size of the (circular) collecting aperture can be exactly calculated as far as

d is concerned. Calling the measured NF MFD, one obtains:

e being the aperture radius on the fibre face. For example, if a 50 //m-core

step-index fibre is employed to collect light on the image plane, this corresponds to

a systematic error of + 0.05 jm (for a 9 yum-MFD). Such an exact calculation is not

possible for the FF MFD d. : then one should have to employ rather lengthy numerical

procedures. A perturbative calculation at 0(e ) [viz. neglecting 0(e ) terms and

) 188
higher ones] for shows that the exact d^. is equal to the measured one, a conclusion

that can be accepted with a sufficient degree of confidence since a similar

perturbative evaluation applied to d^ at the same order gives the same result of

Bq.(3).

3. DIFFRACTION AND NUMERICAL PROCESSING EFFECTS

Diffraction effects due to the finite aperture of the INF lens (assumed to have a

NA = 0.65) have been taken into account calculating the image TNFI distribution by

means of a diffraction integral over the system finite pupil. Noise effects have been

simulated superposing to the image TNFI signal a random contribution according to the

assumed dynamic range of the INF set-up (40 dB). Then we have numerically calculated

the NF and FF MFDs according to Eq.(l) and (2); suitable noise filtering techniques

have been employed in performing numerical derivatives on "measured" quantities. To

start the investigation, we have considered step index fibres, the simplest case. A

first structure (representing a conventional fibre dispersion optimized § 1300 nm) has

a core diameter of 8 yum, a NA of 0.10 and theoretical MFDs 6^=13.45 and d

f

=12.27 /vm at

1550 nm. We found a systematic error of + 0.06 //m on d^ (due to the combined effects

of diffraction and of the smoothing procedures enployed) and of + 0.05 yum on d^

(similarly due to diffraction and numerical processing); the corresponding random

contributions (due to noise) are + 0.02 yum and less than 0.01 yt/m respectively. A

second step structure with core radius of 4.4 yi/m and NA = 0.18, whose theoretical MFDs

at the same wavelength as before are respectively 4^=7.51 and d^=6.83 tm, has been

considered as an example of a dispersion-shifted fibre. In this case the error on d^

is given by a systematic contribution of + 0.03 ytvm and a random term of + 0.01 /urn, while for d^ we have a + 0.09 //m systematic error and a random error lower than 0.01

//m. Finally, we considered a multi-cladding dispersion-flattened fibre (of total

radius 7 (M and NA 0.135) with d =8.54 and d =7.20 yt/m @ 1550 nm. Now d^ is affected

by a systematic error of - 0.02 yum and by a sensible random contribution of + 0.17 /vm.

Also in this case the MFD d^ is more stable, being affected by a systematic error of

+ 0.03 yum and a random term of + 0.03 ywm.

4. COLLUSIONS

Taking into account the error causes discussed in Sections 2 and 3, the total MFD

error for the considered structures is reported in Table I, Three kinds of systematic

terms have been pointed out. A first one depends on the finite size of the detector.

In practical cases this contribution is slightly less than 1%. A second systematic

term is due to diffraction effects. Diffraction causes the presence of Fresnel rings

189
in the image NF, leading to an overestimation or an underestimation of the MFDs

depending on the integration interval determined by the dynamic range and the lens NA

of the set-^up. Finally, processing techniques reduce noise fluctuations but can also

reduce the information content about the intensity distribution. Operating at the

best trade-off point between these contrasting effects, this leads, in general, to

overestimate and d^ of about 1%. In summary, after the elimination of the

systematic errors, we can assess that the MFD determination accuracy from TNFI

distribution measurements is within +0.05 //m for d^ and + 0.03 jjm for d^ in

single-step structures. In more complicated refractive-index distributions (DF SMF)

noise effects reduce the accuracy on d^ to *» + 0.20 /um, and on d^ to + 0.06 /m, a less

dramatic increase. In fact, the FF MFD d

f

is less sensitive to these fluctuations

because the smoothing procedures and noise filtering techniques are employed in Bq.(2)

on both E(r) values and on E'(r) ones; for d^, on the contrary, noise fluctuations,

though smoothed out, are weighted by the factor r l(see Eq.(l)) and then are more

effective. Surprisingly we have found that in TNFI measurements the accuracy on d^ is

better than that on d^. This fact confirms the validity of the CCITT reccommendation

individuating in d^ the more convenient MFD definition to be used in SMF

characterization.

REFERENCES

1 - G. Coppa, P. Di Vita, U. Rossi, El. Lett., Vol.19 (1983) pp.293-294.

2 - C. Pask, Electron. Lett., Vol. 20 (1984), pp. 144-145.

Table I - Total error on MFDs

profile MFD random systematic

conv. step-i.

ii it + 0.05 /urn ± 0.03 /urn + 0.09 /m.

+ 0.05 jjm

DS step-i

± 0.04 fjm

± 0.03 /m

+ 0.08 fm

+ 0.09 //m it ii

DF multi-step + 0.20 jm

± 0.06 im

+ 0.03 fm

+ 0.03 jm ti it

190
INDEX OF AUTHORS

T.F. Adda

G.M. Alameel

0. Aim

N . Amano

W.T. Anderson

M. Artiglla

J.G.N. Baines

P. Beaud

A.F. Benner

G. Beresford

L. Bjerkan

R.J. Black

H. M. Blount

M.A. Brandtner

E.L. Buckland

J . Bures

L.E. Busse

25

25

89

171

131

187

153,165

7

43

175

89

139

107

161

15

139

79

E. D. Head

R.K. Hickernell

D.R. Hjelme

W. Hodel

G.E. Holmberg

A. Hordvik

T. Horiguchi

Y. Horiuchi

I. P. Januar

R.M. Jopson

A.F. Judy

R.M. Kanen

F. P. Kapron

J. P. Kilmer

K. Kitagawa

K.W. Kowaliuk

157

65

103

7

107

85

1

11

103

25

25

131

35

131

11

123

R. Calvani

R. Caponi

A.H. Cherin

F. Cisternino

G. Coppa

L. Curtis

S.K. Das

C. DeBernardi

F.V. DiMarcello

P. Di Vita

T.J. Drapela

C. K. Eoll

M. Eriksrud

J . Ferner

D. A. Fishman

D.L. Franzen

53

53

157

53

187

39,135

25

61

69

187

179

161

85

123

43

157

S. Lacroix

D.R. Larson

S . Lauritzen

P.J. Lemaire

R.S. Lowe

D . Marcus - Hanks

M. Meliga

A.R. Mickelson

S. Morasca

H. Nagai

Y. Namihira

E. Nesset

D.H. Nettleton

M. Nishimura

J. Noda

K. Noda

R.P. Novak

15

139

65

85

69

93

175

61

103

61

19

11

89

115

127

19

171

7

A.O. Garg

H.H. Gilgen

F. Gonthier

T.J. Gryk

M.J. Hacker

D.A. Hall

A.G. Hallam

W.H. Hatton

161

m

139

107

111,157

143

153

127

C. N. Okafor

T. Ooka

T. Oshimi

J.R. Parton

Z. Pasturczyk

D. W. Peckham

J.W. Peters

R.J. Phelan, Jr

119

11

11

183

93

119

35

65

191
CD. Poole 47

M. Potenza 187

H.M. Presby 43

D.E. Put land 165

K.W. Raine 153,165

P.R. Reitz 31

R.P. Salathe 7

C. Saravanos 93

M.J. Saunders 149

J. Schutz 7

V. Shah 39,135

J.R. Simpson 69

R.W. Smith 143

B. Sordo 61

D . A. Svendsen 175

K. Takada 19

IN • X d.J\d LU 1 9

M. Tateda 1

E . A . Thomas 35

U . B . Unrau 97

I . P . Vayshenker 103

H . Wakabayashi 11

G.R. Walker 57

N.G. Walker 57

H.P. Weber 7

S. Yang 103

M . Young 157

W.C. Young 39,135

192
U.S. DEPT. OF COMM. 1. PUBLICATION OR 2. Performing Organ. Report No. 3. Publication Date

BIBLIOGRAPHIC DATA

REPORT NO.

September 1988 SHEET (See instructions) NBS/SP-748

4. TITLE AND SUBTITLE

Technical Digest, Symposium on Optical Fiber Measurements, 1988

5. AUTHOR(S)

G. W. Day and D. L. Franzen, Editors

6. PERFORMING ORGANIZATION (If joint or other than NBS. 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

Previous Symposia were held in 1980, 1982, 1984, and 1986. Technical Digests

of those were published as NBS Special Publications 597, 641, 683, and 720.

PJ Document describes a computer program; SF-185, FlPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here)

This digest contains summaries of 42 papers presented at the Symposium on

Optical Fiber Measurements, held September 20-21, 1988, 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

13. AVAILABILITY

f~X~] Unlimited

| | For Official Distribution. Do Not Release to NTIS

|"X~1 Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C.

20402.

[X] Order From National Technical Information Service (NTIS), Springfield, VA. 22161

14. NO. OF

PRINTED PAGES

202

15. Price

U.S. GOVERNMENT PRINTING OFF ICE: 1988- 576-21 2/ 851 97 USCOMM-DC 0043-P80



Technical Publications

Periodical

Journal of Research—The Journal of Research of the National Bureau of Standards reports NBS research

and development in those disciplines of the physical and engineering sciences in which the Bureau is active.

These include physics, chemistry, engineering, mathematics, and computer sciences. Papers cover a broad

range of subjects, with major emphasis on measurement methodology and the basic technology underlying

standardization. Also included from time to time are survey articles on topics closely related to the Bureau's

technical and scientific programs. Issued six times a year.

Nonperiodicals

Monographs—Major contributions to the technical literature on various subjects related to the Bureau's scientific and technical activities. Handbooks—Recommended codes of engineering and industrial practice (including safety codes) developed in cooperation with interested industries, professional organizations, and regulatory bodies.

Special Publications—Include proceedings of conferences sponsored by NBS, NBS annual reports, and other

special publications appropriate to this grouping such as wall charts, pocket cards, and bibliographies.

Applied Mathematics Series—Mathematical tables, manuals, and studies of special interest to physicists,

engineers, chemists, biologists, mathematicians, computer programmers, and others engaged in scientific and

technical work.

National Standard Reference Data Series—Provides quantitative data on the physical and chemical properties

of materials, compiled from the world's literature and critically evaluated. Developed under a worldwide program coordinated by NBS under the authority of the National Standard Data Act (Public Law 90-396).

NOTE: The Journal of Physical and Chemical Reference Data (JPCRD) is published quarterly for NBS by

the American Chemical Society (ACS) and the American Institute of Physics (AIP). Subscriptions, reprints,

and supplements are available from ACS, 1155 Sixteenth St., NW, Washington, DC 20056.

Building Science Series—Disseminates technical information developed at the Bureau on building materials,

components, systems, and whole structures. The series presents research results, test methods, and perfor- mance criteria related to the structural and environmental functions and the durability and safety

characteristics of building elements and systems.

Technical Notes—Studies or reports which are complete in themselves but restrictive in their treatment of a

subject. Analogous to monographs but not so comprehensive in scope or definitive in treatment of the subject

area. Often serve as a vehicle for final reports of work performed at NBS under the sponsorship of other

government agencies.

Voluntary Product Standards—Developed under procedures published by the Department of Commerce in Part 10, Title 15, of the Code of Federal Regulations. The standards establish nationally recognized re- quirements for products, and provide all concerned interests with a basis for common understanding of the

characteristics of the products. NBS administers this program as a supplement to the activities of the private sector standardizing organizations.

Consumer Information Series—Practical information, based on NBS research and experience, covering areas of interest to the consumer. Easily understandable language and illustrations provide useful background

knowledge for shopping in today's technological marketplace.

Order the above NBS publications from: Superintendent of Documents, Government Printing Office,

Washington, DC 20402.

Order the following NBS publications—F1PS and NBSIR 's—from the National Technical Information Service, Springfield, VA 22161.

Federal Information Processing Standards Publications (FIPS PUB)—Publications in this series collectively

constitute the Federal Information Processing Standards Register. The Register serves as the official source of

information in the Federal Government regarding standards issued by NBS pursuant to the Federal Property

and Administrative Services Act of 1949 as amended, Public Law 89-306 (79 Stat. 1127), and as implemented

by Executive Order 1 1717 (38 FR 12315, dated May 11, 1973) and Part 6 of Title 15 CFR (Code of Federal

Regulations).

NBS Interagency Reports (NBSIR)—A special series of interim or final reports on work performed by NBS

for outside sponsors (both government and non-government). In general, initial distribution is handled by the

sponsor; public distribution is by the National Technical Information Service, Springfield, VA 22161, in paper

copy or microfiche form.
U.S. Department of Commerce

National Bureau of Standards

Gaithersburg, MD 20899

Official Business

Penalty for Private Use $300